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TL 1 


G62 


.A3 


no . 


FH WA- 


RD- 


77-1111 


c.2 J 



brt No. FHWA-RD-77-111 



fIGN OF ZERO -MAINTENANCE PLAIN 
JOINTED CONCRETE PAVEMENT 



Vol. I -Development of Design Procedures 




s r*ns o* 



June 1977 
Final Report 



Dept. of Transportation 



DEC 21 1978 



Library 



1 



This document is available to the public 
through the National Technical Information 
Service, Springfield, Virginia 22161 



Prepared for 

FEDERAL HIGHWAY ADMINISTRATION 
Offices of Research & Development 
Washington, D.C. 20590 



FOREWORD 



Volume I presents the development of comprehensive procedures for the structural 
design of "Zero-Maintenance" plain jointed concrete pavements for heavily 
trafficked roadways. These design procedures are based upon results from long- 
term field studies, comprehensive mechanistic analysis, and laboratory studies. 

This report completes a set of three prepared by the University of Illinois 
under research contract with the Structures and Applied Mechanics Division, 
Office of Research of the Fed ^al Highway Administration. The first report 
is FHWA-RD-76-105, "Zero-Maintenance Pavements: Results of Field Studies on 
the Performance Requirements and Capabilities of Conventional Pavement Systems." 

Procedures and specifications for designing the concrete slab, subbase, shoulders, 
joints, and subsurface drainage are presented in FHWA-RD-77-112, Volume II, 
Design Manual . 

The report is intended primarily for research and development audiences. Copies 
are being distributed accordingly by transmittal memorandum. 



.heWey 



Charles F. SchePfey 
Director, Office of Research 
Federal Highway Administration 



NOTICE 



This document is disseminated under the sponsorship of the Department of 
Transportation in the interest of information exchange. The United States 
Government assumes no liability for its contents or use thereof. 

The contents of this report reflect the views of the authors who are 
responsible for the facts and the accuracy of the data presented herein. 
The contents do not necessarily reflect the official views or policy of the 
Department of Transportation. 

This report does not constitute a standard, specification or regulation. 

The United States Government does not endorse products or manufacturers. 
Trademarks or manufacturers' names appear herein only because they are 
considered essential to the object of this document. 



Technical Report Documentation Page 



1. Report No. 

FHWA-RD-77-111 



2. Government Accession No. 



3. Recipient's Catalog No. 



4. Title and Subtitle 



Design of Zero-Maintenance Plain Jointed Concrete 
Pavement, Vol. I - Development of Design 
Procedures 



5. Report Date 

8 June 1977 



6. Performing Organization Code 



7. Author's) 

Michael I. Darter 



8. Performing Organization Report No. 



FHWA-RD-77-111 



9. Performing Organization Name and Address 

Department of Civil Engineering 
University of Illinois at Urbana-Champaign 
Urbana, Illinois 61801 



10. Work Unit No. (TRAIS) 



" I) 



11. Contract -or Grant No. 

DOT-FH-11-8474 



12. Sponsoring Agency Name and Address 

Federal Highway Administration 
U.S. Department of Transportation 
Office of Research 
Washington, D.C. 20590 



13. Type of Report and Period Covered 

Final, 1975-1976 



14. Sponsoring Agency Code 



87 



15. Supplementary Notes 



FHWA Project Monitors 



Floyd J. Stanek, Contract Manager 
Thomas J. Pasko 
William J. Kenis 



16. Abstract 



Comprehensive procedures for the structural design of "zero-maintenance" 
plain jointed concrete pavements for heavily trafficked roadways are developed, 
The term "zero-maintenance" refers to the structural adequacy of the pavement 
lanes and shoulder. Thus, a "zero-maintenance" pavement would not require 
maintenance such as patching, joint repair, crack repair, grinding, and 
overlays. The design procedures are based upon results from long-term field 
studies, comprehensive mechanistic analyses, and laboratory studies. Both a 
serviceability-performance analysis and a concrete fatigue analysis are used 
in the structural design. Procedures are developed for designing the concrete 
slab, subbase, shoulders, joints and subsurface drainage. Example designs 
are included with sensitivity and incremental cost analyses. A r.nmnut.er 
program, JCP-1 , was written to assist in the design calculati 
manual (Vol. II) was developed based upon the results documen 




17. KeyWords 

Pavement, design, concrete, fatigue, 
serviceability, performance, 
maintenance, distress 



18. Distribution Statement 



No restrictions. This document is 
available to the public through the 
National Technical Information 
Service, Springfield, Virginia 22161 



19. Security Classif. (of this report) 

Unclassified 



20. Security Classif. (of this page) 

Unclassified 



21. No. rf Pages 

261 



22. Price 



Form DOT F 1700.7 (8-72) 



Reproduction of completed page authorized 



PREFACE 

"Design of Zero-Maintenance Plain Jointed Concrete Pavement, Vol. I - 
Development of Design Procedures" provides documentation of the development 
of procedures for the design of heavily trafficked highway pavements. The 
objective of the design is to provide pavements which will perform rela- 
tively maintenance-free over a selected design period. The term "zero- 
maintenance" refers only to structural maintenance such as patching, crack 
filling, slab replacement, and overlay. Procedures are developed for designing 
the following components of plain jointed concrete pavements: Portland 
cement concrete slab, subbase, shoulders, joints, and subsurface drainage. 
A computer program, called JCP-1 , is used to provide serviceability/performance 
and fatigue damage data for structural design of the pavement. Manual pro- 
cedures are also included to structurally design the pavement based on 
serviceabil ity/performance. 

This study was conducted at the Department of Civil Engineering, 
University of Illinois at Urbana-Champaign under sponsorship of the U.S. 
Department of Transportation, Federal Highway Administration. The princi- 
pal investigators of the study are Dr. Michael I. Darter and Dr. Ernest 
J. Barenberg. The author wishes to sincerely thank the several persons who 
contributed directly to the development of this manual, including: Mr. 
Jihad Sawan, Miss H. S. Yuan, Mr. Amir M. Tabatabaie, Mr. Clive Campbell, 
Professor Marshall R. Thompson, and Professor Barry J. Dempsey. Thanks 
are also due to numerous state highway engineers from many states for 
providing considerable data and other assistance. Thanks are also due the 
FHWA project monitors Mr. William J. Kenis, Mr. Thomas Pasko, and 
Dr. Floyd Stanek, for their assistance and encouragement throughout the 
study. A special note of thanks to Mrs. Karon Webb for typing and editing 

this manuscript. 

i i 



TABLE OF CONTENTS 

LIST OF ABBREVIATIONS 

CONVERSION FACTORS, U.S. CUSTOMARY TO METRIC (SI) UNITS OF 
MEASUREMENT 

CHAPTER 1 INTRODUCTION 1 

1.1 Background 1 

1.2 Research Approach 3 

1.3 General Desiqn Approach 5 

1.4 Limitations 7 

CHAPTER 2 FIELD SURVEY AND DISTRESS 10 

2.1 Field Survey 10 

2.2 Analysis of Distress 19 

CHAPTER 3 STRUCTURAL DESIGN BASED ON SERVICEABILITY/PERFORMANCE ... 35 

3.1 Introduction 35 

3.2 Original AASHO Model 36 

3.3 Development of New Equations 41 

3.3.1 Modified AASHO Equation 41 

3.3.2 New Approach 44 

3.4 Climatic Regional Factor 52 

3.5 Terminal Serviceability Index for Zero-Maintenance Design ... 54 

3.6 Sensitivity of New Performance Equation 56 

CHAPTER 4 STRUCTURAL DESIGN BASED ON PCC FATIGUE 66 

4.1 Finite Element Model 66 

4.1.1 Description of Finite Element Method 67 

4.1.2 Transverse and Longitudinal Joints 68 

iii 



4.1.3 Computer Program 69 

4.1.4 Comparison of Measured and Computed Load Stress .... 70 

4.1.5 Comparison of Computed and Measured Thermal 

Curl Stress 77 

4.2 Effects of Pavement Factors 81 

4.3 Critical Fatigue Location in Slab 101 

4.2.1 Initiation of Cracking - Field Results 101 

4.2.2 Initiation of Cracking - Fatigue Analysis ....... 105 

4.4 Effect of Joint Spacing on Cracking 121 

4.5 Development of Fatigue Damage Analysis 132 

4.5.1 PCC Fatigue 132 

4.5.2 PCC Strength Increase 142 

4.5.3 Lateral Truck Distribution 143 

4.5.4 Thermal Gradients 147 

4.5.5 PCC Fatigue Computation 148 

4.6 Limiting Fatigue Consumption 156 

CHAPTER 5 DESIGN OF JOINTS, SHOULDERS, AND SUBSURFACE DRAINAGE .... 165 

5.1 Joints 165 

5.1.1 Joint Faulting 165 

5.1.2 Joint Sealant Damage 182 

5.1.3 Transverse Joint Spacing 186 

5.1.4 Joint Load Transfer Device 188 

5.2 Shoulders 197 

5.2.1 PCC Shoulder Design 199 

5.2.2 Asphalt Concrete Shoulder Design 199 

5.3 Subsurface Drainage 202 



IV 



CHAPTER 6 VERIFICATION OF DESIGN 205 

6.1 Design Approach and Computer Program 205 

6.2 Structural Design Verification 208 

CHAPTER 7 CONCLUSIONS AND RECOMMENDATIONS 214 

7.1 Conclusions 214 

7.2 Recommendations 216 

REFERENCES 217 

APPENDIX COMPUTER PROGRAM JCP-1 FOR ZERO-MAINTENANCE DESIGN 224 

A.l JCP Input Guide 224 

A. 2 Sample Input 231 

A. 3 Flow Chart of Program 233 

A. 4 JCP-1 Program Listing 237 



LIST OF ABBREVIATIONS 

Pavement Section 

AC Asphalt Concrete 

ADT Two-directional Average Daily Traffic 

ATB Asphalt Treated Base 

CJS Contraction Joint Spacing 

CTB Cement Treated Base 

EJS Expansion Joint Spacing 

ESAL Equivalent Single Axle Loads 

GR.B. Granular Base of Subbase 

JCP Jointed Concrete Pavement (Non-reinforced) 

JRCP Jointed Reinforced Concrete Pavement 

LTD Load Transfer Device 

PCC Portland Cement Concrete 

SM Select Material 

Maintenance 
CF Crack Filling 

JF Joint Filling 

P Patching with AC or PCC 

Distresses 
JF Joint Faulting 

JS Joint Spall ing 

CF Crack Faulting 

CS Crack Spall ing 



VI 



i 



LIST OF ABBREVIATIONS (Continued) 

Distresses (Continued) 

TC Transverse Cracking 

B Blowups 

CC Corner Cracking 

LC Longitudinal Cracking 

D "D" Cracking 

PU Pumping 

DC Diagonal Cracking 

Climatic Region 
WF Wet-freeze Region 

W Wet-nonfreeze Region 

DF Dry-freeze Region 

D Dry-nonfreeze Region 



vn 



CONVERSION FACTORS, U. S. CUSTOMARY TO METRIC (SI) 
UNITS OF MEASUREMENT 



U. S. customary units of measurement used in this report can be converted 
to metric (SI) units as follows: 



MULTIPLY 
inches 
feet 

square inches 
square yards 
knots 
pounds 
kips 

pounds per cubic foot 
pounds 
kips 

pounds per square inch 
pounds per cubic inch 
gallons fU. S. liquid) 
Fahrenheit degrees 



BY TO OBTAIN 

2.54 centimeters 

0.3048 meters 

6.4516 square centimeters 

0.83612736 square meters 

0.5144444 meters per second 

0.45359237 kilograms 

0.45359237 metric tons 

16.018489 kilograms per cubic meter 

4.448222 newtons 

4.448222 kilonewtons (kN) 

6.894757 kilopascals 

2.7144712 kilopascals per centimeters 

3.785412 cubic decimeters 

5/9 Celsius degrees of Kelvins* 



* To obtain Celsius (C) temperature readings from Fahrenheit (F) readings, 
use the following formula: C = (5/9)(F-32). To obtain Kelvin (K) 
readings, use: K = (5/9)(F-32) + 273.15. 



vm 



CHAPTER 1 
INTRODUCTION 

This report describes the development of comprehensive procedures for 
the design of "zero-maintenance" plain jointed Portland cement concrete 
pavement. Based upon the work described herein, a design manual, Volume II (Ref. 1), 
was written which gives step by step procedures for designing zero-maintenance 
plain jointed concrete pavements (JCP). The term "zero-maintenance" refers 
to the structural adequacy of the pavement travel lanes and shoulders. Thus, 
a zero-maintenance pavement would not require maintenance such as: patching, 
joint repair, crack repair, grinding, and overlays. However, activities 
such as mowing, guard rail repair, stripping, providing skid resistance, 
wear from studded tires, geometric obsolescence, and subsequent widening 
to increase capacity are not included in the definition of maintenance in 
this report. 

1.1 BACKGROUND 

Many highways in urban and suburban areas are being subjected to heavy 
traffic volumes which cause rapid deterioration and premature failure of 
pavements. Hence, considerable maintenance is required, but scheduling of 
remedial and preventative maintenance is almost impossible without closing 
lanes and producing massive traffic jams, accidents, and delays to the 
traveling public. Often routine maintenance is completely neglected, thus 
causing even more accelerated deterioration of pavements. When maintenance 
is performed, it is usually during off-hours or at nights. Under such 
conditions, repairs are often rushed and/or performed with inadequate equip- 
ment and with inadequate room to manuever. The repairs are often inefficiently 

1 



done because of logistics, traffic interference, and workers toiling under 
hazardous conditions. The cost of traffic control is a major item in any 
maintenance budget, but especially under the conditions described, and the 
cost of delays to the motorist because of lane closure or detours from the 
expressway for maintenance operations accumulates at a fantastic rate. 

The fundamental question which underlies this research is: "how to 
design and build conventional pavements (or optimized conventional pavements 
in which inherent weaknesses are eliminated) to serve exceptionally heavy 
traffic without requiring maintenance and providing satisfactory ' rideability' 
for twenty years, and with only routine maintenance for an additional 10 
years?" Initially five types of conventional pavements were considered 
including: jointed reinforced concrete, continuously reinforced concrete, 
flexible (including thick asphalt layers and full depth asphalt), composite 
(asphalt over concrete), and jointed plain concrete pavements. The first 
phase of the study included extensive field surveys and evaluations of 
over 70 heavily trafficked pavements, and the following was determined: 

1. Types and causes of distress and maintenance applied on heavily 
trafficked pavements; 

2. Adequacy of commonly used design procedures to obtain maintenance- 
free pavements; 

3. Limiting criteria for use in designing maintenance-free pavements; 
and 

4. Maximum maintenance- free lives of conventional pavements. 

These results were documented in a report bv Darter and Barenberg (Ref. 2). 
The second phase of the study involved the development of zero-maintenance 
design procedures for jointed plain concrete pavements which is contained in 
this report. 



1.2 RESEARCH APPROACH 

The research approach used to develop the design procedures 
is illustrated in Figure 1.1. Field studies were conducted and 
plain joined concrete pavements were examined in 10 highway agencies and 
extensive data collected. The types, causes, and ways to eliminate or 
minimize the significant distresses were identified based upon the experience 
of local pavement engineers and project staff, previous research studies, 
and analytical studies conducted as part of the project. Existing design 
procedures were critically evaluated as to their ability to provide zero- 
maintenance pavements and their limitations determined. Limiting criteria 
were determined for zero-maintenance design (including terminal serviceability 
and allowable fatigue consumption). All available long term performance data 
of plain jointed concrete pavements were complied which included 25 sections 
from the original AASHO Road Test that have been under regular traffic since 
1962 on 1-80 in Illinois and 12 other projects located in various climatic 
regions which vary in age from 6 to 34 years. Analytical models and proce- 
dures for slab stress/strain computation and fatigue damage were developed, 
and a new serviceability/performance model was derived. A comprehensive 
fatigue analysis procedure was developed and verified that gives accumu- 
lated fatigue damage at the most critical point in the slab considering 
both traffic load applications and curling of the slab. A comprehensive 
yet practical design procedure was developed that considers both fatigue 
damage and serviceability loss in selection of the final pavement structure. 
Design recommentations were also developed for other components of the pave- 
ment system, including shoulders, joints, subbase, and subsurface drainage 
based upon results from the overall study and other research results. 



Evaluate 
Existing 
Design 
Procedures 



Field Studies 

Pvt. Surveys 
Interviews 
Data Collection 



I 



Identify Types, 
Causes, Ways To 
Eliminate Distress 



Select Data Base 
And Analytical Models 



Select 
Limiting 
Criteria For 
Design 



Develop Comprehensive 
Design Procedures 



Verify Procedures 



Figure 1.1. Research Approach to Develop Zero-Maintenance 

Design Procedures for Plain Jointed Concrete Pavements 



1.3 GENERAL DESIGN APPROACH 

The general design approach consists of (1) determination of material 
properties and structural thicknesses of the PCC slab and the subbase, 

(2) selection of joint spacing, configuration, load transfer, and sealant, 

(3) determination of shoulder type and dimensions, and (4) subsurface drain- 
age provisions. These components are designed as a system to ensure com- 
patibility. A flow diagram showing the major design steps is shown in 
Figure 1.2. 

The structural design procedures consist of both a slab fatigue analysis 
and also a serviceability/performance analysis. The final structure design 
is based upon both of these considerations to ensure more comprehensive 
analysis of pavement performance. A computer program is included that 
provides fatigue damage and serviceability/performance data used for selec- 
tion of the structural design. The program is named JCP-1 and is written 
in FORTRAN. A manual procedure is also included to determine structural 
design based on serviceability/performance. 

The procedure shown in Figure 1.2 is iterative, indicating that there 
are, of course, more than one zero-maintenance design alternative. The 
design that gives the minimum construction cost is generally selected as 
the optimum design as long as it meets all of the limiting design criteria. 

The justification for construction of a zero-maintenance design is 
based upon an economic analysis. The increased costs to construct a zero- 
maintenance pavement over that of a conventional pavement must be compared 
with the costs resulting from maintenance, rehabilitation, and user delay 
if a conventional pavement is constructed. These costs must be computed over 
a given analysis period such as 20 years. Procedures have been developed by 



CD 

> 



O 

c 



C 

in 
Q 



O 




i 



Conduct 

Serviceability 

Analysis 



EZ 



Select 
Structural 
Thickness 
Materials 



And 



Joint Design ', 
Spacing, Dimen. 
Sealant, Load Tran 



Shoulder 
Design 



Select 
Overall 
Pavement 
Design 



Subsurface 
Drainage 



Conduct Econ. 
Analysis-Select 
Optimum Design 



Figure 1.2. Zero-Maintenance Design Procedure for 
Plain Jointed Concrete Pavements. 



Butler (Ref. 3) for FHWA to estimate the maintenance, rehabilitation, and 
user delay costs of conventional pavements. 

1.4 LIMITATIONS 

An important question that was posed many times during the develop- 
ment of these procedures is: Can a pavement be constructed that actually 
lasts 20 or more years without requiring structural maintenance such as 
crack repair, overlay, grinding, joint repair, patching, etc.? The field 
survey revealed that there are several plain jointed concrete and other 
pavement types that have performed maintenance-free for 15 to 27 years under 
heavy traffic. Therefore, it is possible to design and construct a pave- 
ment with this performance requirement. It requires, however, a most compre- 
hensive and thorough design approach that considers all significant details 
to tailor the design to local conditions unique to the project. Although 
detailed recommendations are provided which are useful to 
most design situations, there is no substitute for engineering experience, 
which in certain instances may overrule specific recommendations given 
herein. 

The design procedures contained herein have been developed using 
the most comprehensive mechanistic models available, and also long term 
measured pavement performance data. They have been verified using all 
data available to the project staff and found to give reasonable results. 
However, there are several aspects that are not as fully reliable as others 
due to lack of data and technology and these limitations must be carefully 
considered. 

1. PCC Durability - The deterioration of PCC from any of several 
causes will cause a reduction of pavement maintenance-free life. Although 

7 



specific recommendations are given to minimize the occurrence, it may not 
be possible in some regions to prevent deterioration with existing materials. 

2. Pavement Growth - The infiltration of a considerable amount of 
incompressibles into joints may result in pavement growth at bridge ends. 
Therefore, high type joint sealants with long performance life must be 
provided to minimize this occurrence. 

3. Joint Faulting - Recommendations herein specify that dowel bars 
must be used in most all pavements, with the possible exception of pave- 
ments with very low truck volumes and pavements located in warm dry climates. 
If dowels are not used in other conditions, joint faulting may occur which 
would reduce the maintenance-free life of the pavement. 

4. Construction - The failure to achieve construction quality as 
required in the specifications may have a serious effect on reducing the 
maintenance-free life of the pavement. A thorough inspection of the pave- 
ment should be conducted after construction to ascertain if any deficiencies 
exist, which would result in a reduced maintenance-free life. These should 
be corrected, if possible. 

5. Design in Various Climates - Results from field surveys indicate 
that plain jointed concrete performs differently in different climates. 
Some climatic effects can be quantified directly herein, but an empirical 
climatic factor is still needed to help adjust for the difference 

in performance. This factor is not sufficiently verified and should be 
adjusted if it does not provide reasonable results in certain climates. 

6. Traffic Estimation - A considerable effort has been made to specify how 
to obtain reasonable traffic estimates for design. The most crucial factor 
being the axle load distribution. The designer must carefully estimate all 

8 



traffic inputs using the best sources of data available. Underestimation 

of traffic to a significant degree may result in a pavement structure 

not capable of lasting throughout the design period in a maintenance-free 

condition. 

The results of this study are presented in the following sequence: 

Chapter 2 - Description of the field survey and distress in JPCP. 

Chapter 3 - Development of a new serviceability-performance design 
model . 

Chapter 4 - Development of a comprehensive fatigue analysis procedure 

Chapter 5 - Design recommendations for joints, shoulders, and sub- 
surface drainage. 

Chapter 6 - Verification of zero-maintenance design procedures. 

Chapter 7 - Conclusions and recommendations for implementation of 
design procedures. 



CHAPTER 2 
FIELD SURVEY AND DISTRESS 

2.1 FIELD SURVEY 

Field surveys of heavily trafficked plain jointed concrete pavements 
were conducted in the U.S. and in one Canadian province. Extensive inter- 
views, condition surveys, and data collection were made in 11 highway agencies 
where 37 projects were selected over a variety of climatic conditions. 
These projects were used for detailed performance evaluation and other 
analyses. The general location of these projects is shown in Figure 2.1. 

Discussions with agency personnel and condition surveys were conducted 
in each area visited as described below: 

1. Detailed discussions were held with administration, design, con- 
struction, and maintenance personnel of the agency. Subjective information 
was thus gathered for pavement performance, types of distress, recommended 
design practices, critical design limits, and construction and maintenance 
practices in the area. Some of this information is contained in a previous 
report (Ref. 2). 

2. Condition surveys were conducted on several projects in each area, 
some of which were selected for inclusion in the project analyses. Attempts 
were made to collect data from each project as follows: 

a. A surface condition survey was conducted over a typical 

2000 ft (610 m) portion of the pavement and all distress was re- 
corded. The outer traffic lane was surveyed in all cases since 
this was the lane having the highest truck volume. 

b. Present Serviceability Rating (PSR) was estimated by the 



10 




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project staff member(s), or when available, the Present Servicea- 
bility Index (PSI) was obtained from the agency. 

c. A general drainage evaluation was made. 

d. The pavement was photographed both with 35 mm camera and with 
a Super-8 movie camera. 

e. The date of opening to traffic and the original pavement cross 
section were obtained from the agency. Any available material 
data for various layers and the subgrade measured during or after 
the construction of the project were also obtained. 

f. Traffic data including ADT, percent trucks, axle load distri- 
butions, lane distributions, directional distribution, and average 
axles per truck were obtained from the agency. 

g. Climatic data were obtained from published Weather Bureau and 
other sources. 

h. Opinions of the local engineers as to the reasons for pavement 

distress, if any, were obtained. 

i". Previous maintenance performed on the section was determined. 

All agencies were yery cooperative and helpful in collecting and pro- 
viding the requested data. A summary of data for all 37 projects is given 
in Tables 2.1 and 2.2. Performance data for each project are contained in 
other chapters. 

3. The following guidelines were used in project selection: 

a. Age - Longevity was important in that generally only projects 
over 10 years in age were selected. 

b. Maintenance - Only projects which had not been overlayed were 



12 



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15 



selected. Past routine maintenance was not a consideration and 
both pavements performing maintenance-free and those receiving 
maintenance were selected. 

c. Traffic - Pavements that were heavily trafficked were given 
preference in selection. All of the projects are heavily 
trafficked urban or rural freeways. 

d. Data Sample - The project sample represents only a small 
portion of all heavily trafficked plain jointed concrete pave- 
ments in the U.S. However, many of the projects are representa- 
tive of others in the general area. JCP-28 in Los Angeles is 
typical of many pavements in that area that have been constructed 
for many years (others in this category include JCP-26, 27, 31, 
34, and 35). The 25 sections in Illinois (JCP-1 to 25) are from 
the original AASHO Road Test that were left inservice on 1-80 
from 1962 to 1974. Three projects represent "one of a kind" 
constructed in a given area such as JCP-36 in Detroit (the 
first freeway constructed in 1942), JCP-32 in New Jersey, and 
JCP-29 near Dallas, Texas. Various plain jointed concrete pro- 
jects in other states were observed including Nebraska, Wyoming, 
Florida, Ohio, and Pennsylvania but no data were collected. 
Available literature on the performance of plain jointed concrete 
pavement was also reviewed. 

Four general climatic regions were selected as defined in Table 2.3 
and selected climatic data for these locations are given in Table 2.4. The 
regions are based on moisture and freeze-thaw considerations. Projects were 
selected within each of the four climatic regions. 



16 



Table 2.3. Definition of the Four General Climatic Regions 



Climatic 
Region 



Annual Precipitation (P) 
and Potential Evapo- 
transpiration (E) 



Frost Heave 
and/or Freeze- 
Thaw Damage 



Wet/Freeze (WF) 



Wet/Non-freeze (W) 



Dry/Freeze (DF) 



P > E or 

P > 30 ins (0.76 m) 

P > E or 

P > 30 ins (0.76 m) 

P < E 



Occurs in pavements 
in region* 



Does not occur in 
pavements in region 



Occurs in pavements 
in region* 



Dry/Non-freeze (D) 



P < E 



Does not occur in 
pavement in region 



*Generally in areas having a mean Freezing Index > (Ref. 64) 



17 



Table 2.4. Annual Precipitation, Evapotranspi ration, and Freezing Index 
of Field Project Regions. 



1 2 3 

Project Annual Precipitation Freezing 

Region Precipitation Minus Evaporation Index 

(ins.) (ins.) 



I. WET/FREEZE AREA 

Ottawa, IL 30 + 4* 700 

Detriot, MI 29 +4 500 

Toronto, 0NT 31 +8 750 

New Brunswick, NJ 39 +10 100 

II. DRY/FREEZE AREA 

Salt Lake City, UT 13 -20** 250 

Denver, CO 17 - 8 300 

III. WET/NO-FREEZE AREA 

Atlanta, GA 43 +12 

Seattle, WA 30 +30 

IV. DRY/NO-FREEZE AREA 

San Francisco, CA 17 -10 

Los Angeles, CA 10 -30 

Phoenix, AZ 5 -65 

Dallas, TX 28 -24 



* plus means precipitation is more than evaporation 
** minus means precipitation is less than evaporation 

1 in. = 2.54 cm 

1. Ref. 63 

2. Ref. 63 

3. Ref. 64 



18 



2.2 ANALYSIS OF DISTRESS 

The development of design procedures to provide zero-maintenance per- 
formance requires the consideration and prevention of all distresses that 
require maintenance. Thus, an identification and study of distresses 
occurring in plain jointed concrete pavement is required. Results from 
the field study and other data were analyzed to determine distresses occur- 
ring in conventional heavily trafficked plain jointed concrete pavements. 
Eight distress types were identified in the 37 pavements surveyed as 
summarized in Table 2.5. The "Distressed/Total" column represents the 
number of pavements containing the indicated distress to the total number 
surveyed (or 37). The column "Maintained/Distressed" indicates the number 
of pavements receiving maintenance for the indicated distress to the total 
number distressed. Thus there were, for example, 12 projects exhibiting 
transverse cracking, and 10 of these projects received maintenance for this 
distress. A summary of distress, performance, and maintenance data for 
each project is given in Table 2.6. 

The following conclusions are based on these data: 
(1) Joint Faulting - nearly half of all projects had greater than 
0.05 in. (1.3 mm) mean joint faulting. Even though only 3 of the 16 re- 
ceived maintenance for this distress, it is considered the most serious 
distress because when faulting develops to a level greater than about 
0.20 in. (5.1 mm) it results in an intolerable ride quality and mainte- 
nance is usually performed. Several pavements located in the area surveyed 
have been overlayed or had fault grinding because of excessive joint 
faulting. This distress was considered the most serious problem by those 
engineers interviewed and must be considered in design. Photos of 



19 



Table 2.5. Summary of Distress Types Occurring on Plain 
Jointed Concrete Pavements. 



Type of Distress Distressed/Total** Mai ntained*/Di stressed 

1. Joint Faulting (>0.05 in.) 16/37 3/16 

2. Transverse Cracking 12/37 10/12 

3. Longitudinal Cracking at 3/37 0/3 

Joint 

4. Corner Cracking 1/37 0/1 

5. "D" Cracking at Joint 16/37 16/16 

6. Joint and Corner Spalling 8/37 5/8*** 

( > 3 in. d i a . ) 

7. Joint Seal Damage 35/37 29/35 

8. Settlement 27/37 3/27 



1 in = 25.4 mm 

* Maintenance applied only to distress indicated. 
** Total number of projects included in field study. 
*** Spalls and maintenance patches are very small. 



20 



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22 



serious joint faulting are shown in Figure 2.2 and 2.3 (further discussion 

is given in Chapter 5 on causes). 

C 

(2) Transverse bracking - occurred on 12 projects, and 10 of these 

received maintenance. Thus, transverse cracking i,s a serious distress 
that usually requires maintenance because of crack spalling and faulting. 
Transverse cracking therefore must be considered in design (additional 
information on causes are given in Chapter 4). Photos of typical transverse 
cracking are shown in Figs. 2.4-2.9. 

(3) Longitudinal Cracking - only occurred on three projects and none 
received maintenance. This crack always occurred at the joint about 1-3 
ft. from the slab edge and only extended about 2 ft., as shown in Fig. 2.10 
(further discussion is given in Chapter 4). The cracks were all hairline 
and did not spall . 

(4) Corner cracking - only one project (without dowels) exhibited this 
distress. This project also had skewed joints. Significant faulting and 
pumping was evident on this project as shown in Fig. 2.11. 

(5) "D" Cracking - only occurred on those projects located at the 
AASHO Road Test site. However, it has been observed to exist throughout 
the midwestern U.S. and in other areas. "D" Cracking usually results in 
joint spalling that was subsequently patched on all 16 pavements. This 
distress must be considered in materials selection and design (See Figs. 
2.12-2.14). 

(6) Joint and Corner Spalling - occurred on 8 of the pavements to a 
very small degree. This was generally the result of "D" cracking as shown 

in Figs. 2.13 and 2.14. It indicates however, that higher quality PCC is needed 



23 



MUMMKattam 




Figure 2.2. Joint Faulting on Plain Jointed Concrete Pavement 

Containing Mo Dowels (JCP-28-CA, approximately 0.2 in.) 

1 in = 25.4 mm 




Figure 2.3. Joint Faulting on Plain Jointed Concrete Pavement 
Containing No Dowels (GA). 



24 






mmrmm 



Figure 2.4. Transverse Crack on Plain Jointed Concrete Pavement 
with 15 ft. Joint Spacing (JCP-28-CA). 



1 ft - 0. 30 m 





Figure 2.5. Transverse Crack on Plain Jointed Concrete Pavement 
with 20 ft. Joint Spacing (GA). 



25 



I 




Figure 2.6. 



Transverse Cracking on Plain Jointed Concrete Pavement 
with 25 ft. Joint Spacing (JCP-36-MI). 



1 ft = 0. 30 m 




-<* 







r 






Figure 2.7. Transverse Cracking on Plain Jointed Concrete Pavement 
with 19 ft. Joint Spacing (JCP-37-0ntario) . 



26 




Figure 2.8. Transverse Cracking on Plain Jointed Concrete Pavement 
with 15 ft. Joint Spacing (JCP-27-CA). 

1 ft = 0. 30 m 




f-W* 




HA 
■■ft" 









*:&& 






v.- 



#"•?•"».'■<»»** 



Figure 2.9. Transverse Cracking on Plain Jointed Concrete Pavement 
with 15 ft. Joint Spacing (JCP-l-IL). 



27 




Figure 2.10. Longitudinal Cracking on Plain Jointed Concrete 
Pavement (JCP-26-WA) . 




■MM) 




*^W 



Figure 2.11. Corner Cracking Plain Jointed Concrete Pavement 
without Dowels at Joint (JCP-34-C0). 



28 



"-/, 



Figure 2.12. "D" Cracking of Plain Jointed Concrete Slab at 
Corner (JCP-22-IL). 



»' — i 




Figure 2.13. Spalling at Joint Caused by "D" Cracking of Plain 
Jointed Concrete Slab (JCP-22-IL). 



29 



•'"£%_, >, ...j.S^Ti *;.*!<?/' 5* ' '*.... '-'i 






Figure 2.14. Spalling at Joint Caused by "D" Cracking of Plain 
Jointed Concrete Slab (JCP-21-IL). 



30 



to minimize this distress. None of the joints showed significant spalling 
where "D" Cracking did not exist. 

(7) Joint Seal Damage - this distress occurred on nearly every project, 
however, only a few were maintained. Sealant damage (infiltration of incom- 
pressibles) is shown on one 9 year old pavement in Fig. 2.15. The stripping 
of sealant can also be seen in Figs. 2.13 and 2.14. This indicates that 
(1) the sealants used are not sufficiently durable, and (2) most agencies 
do not maintain joints on heavily traveled routes. The longitudinal joint 
between the PCC traffic lane and asphalt shoulder was usually always open 

and not sealed and water could freely enter most of the joints. 

(8) Settlement - this distress is difficult to observe but did 

■occur at least to some degree on several projects (See Fig. 2.9 for example). 
Its effect is perhaps most obvious on the thicker slab sections at the AASHO 
Road Test where no cracking, faulting, or any significant distress of any 
kind occurred, and yet the serviceability index dropped from about 4.5 to 
3.5 due to increased roughness, apparently from non-uniform slab settlement. 
This indicates the importance of a stable foundation which must be considered 
in design. 

These distresses are all associated with the traveled lanes. The 
shoulders of most of these pavements showed considerable distress (alliga- 
tor and linear cracking, settlement, heaving, longitudinal joint widening 
and spalling, as illustrated in Figs. 2.16-2.17 etc.). Considerable 
maintenance was applied to the shoulders in several projects. Also, this 
distress appears to have increased the distress occurring in the traveled 
lane (i.e. faulting, pumping, and cracking). Hence, the shoulder must also 
be designed for maintenance- free performance to avoid lane closure due to 
shoulder maintenance. / 



31 




■*■■':' 



Figure 2.15. Joint Sealant Damage (Stripping and Infiltration of 

Incompressibles) for 9 Year Heavily Trafficked Freeway 
(JCP-30UT). 



32 




Figure 2.16. Settlement and Spalling of Shoulder (JCP-29-TX) 





Figure 2.17. Alligator Cracking in Shoulder near Longitudinal 
Joint (JCP-30-UT). 



33 



In summary, the following distress types commonly occurring in heavily 
trafficked plain jointed concrete pavements must be considered in design and 
thereby prevented to provide a zero-maintenance pavement: (1) joint faulting, 
(2) transverse cracking, (3) "D" cracking, (4) Joint and corner spalling, 
(5) joint seal damage, (6) settlement of the slab, and (7) shoulder distress. 
There are obviously other distresses that could occur (i.e. scalling, 
blowups, longitudinal cracking, etc.), however, these seven are considered 
the major types presently occurring on conventional plain jointed concrete 
pavements. 



34 



CHAPTER 3 
STRUCTURAL DESIGN BASED ON SERVICEABILITY/PERFORMANCE 

3.1 INTRODUCTION 

Structural design based upon serviceability/performance permits the 
general consideration of several distress types since the serviceability 
index is a function of roughness and distress. The original development of 
the serviceability index correlated it with roughness, cracking and patching 
(Ref. 46), and other studies have correlated it with only roughness (Ref. 65) 
and only distress (Ref. 2). Pavements designed based on serviceability/ 
performance would thus generally consider the following distresses: joint 
faulting, cracking, joint spalling, and differential settlement of slabs. 

The two year AASHO Road Test provided data that were used to derive 
empirical equations that expressed the relationship between the number of 
traffic load applications and magnitude, the serviceability index (SI), 
and slab thickness. This equation was derived by the Road Test staff, and 
extended by the AASHO Committee on Design of Rigid Pavements (Ref. 60), 
and has been used extensively in concrete pavement design since 1962. A 
detailed evaluation of this design equation was conducted by Darter and 
Barenberg (Ref. 2), and several deficiencies were noted. When the equation 
was used to "redesign" pavements that had provided maintenance-free perfor- 
mance, the results showed that it usually provided inadequate structures. 

One of the major deficiencies of the AASHO performance equation is 
that is is based on only two years data. Since the end of the AASHO Road 
Test in 1960 the Illinois Department of Transportation (IDOT) has monitored 
the performance of 25 of the original sections of the Road Test after 
they were opened to traffic as part of 1-80 in 1962. A summary of 25 



35 



plain jointed concrete pavement sections and their performance that were 
monjtored from 1962 to 1974 are given in Table 3.1. Slab thickness ranges 
from 8 to 12.5 inches (203-318 mm). This data, along with the estimated 
1 8- kip (80 kN) equivalent single axle load (ESAL) applications, were used 
to derive a new performance equation. The data represents 16 years of time 
and nearly 20 million applied 18-kip (80 kN) (ESAL) applications. 

New performance equations were developed using both the same approach 
applied by the Road Test staff, and a new approach found necessary to 
provide a more accurate representation of the relationship between the 
serviceability index, accumulated 18-kip (80 kN) ESALs, and slab thickness. 
Data was also obtained from 12 other plain jointed concrete projects located 
in other regions of the U.S. to provide data on the regional or climatic 
effect on performance. The development of the original AASH0 equation 
is first described, then the development of the new equations based upon 
the 16 year data is given. 

3.2 ORIGINAL AASHO MODEL 

The Road Test provided data to derive empirical relationships between 
PCC slab thickness, load magnitude, axle type, number of load applications, 
and serviceability index of the pavement for Road Test conditions (i.e., 
specific environment and materials) using multiple regression analyses. 

log 1Q W = log P + G/B (3.1) 

where 

W = axle load applications, for load magnitude LI and axle type 
load L2, to a serviceability index of P2. 

log p = 5.85 + 7.35 log(H + 1) + 4.62 log(Ll + L2) + 3.28 log(L2) 



36 



Table 3.1. Summary Performance Data for 25 Sections of Plain Jointed 
Concrete Pavements from AASHO Road Test That Are Inservice 
on 1-80 (1958-1974).* 





Slab 
Thickness (in ) 


Subbase 
Thickness (in ) 






SI 


Data 






Section No 


1962 I 


1968 


1969 


1971 


1972 


J974 


672 


8.0 


3 


4.10 


3.44 


3.21 


2.61 


2.76 


1.69 


658 


8.0 


6 


4.10 


3.14 


2.83 


2.61 


2.11 


1.69 


652 


8.0 


9 


4.10 


3.01 


3.02 


2.98 


2.57 


1.69 


552 


9.5 





4.30 


3.36 


3.52 


3.14 


3.14 


2.87 


676 


9.5 


3 


4.00 


3.21 


3.21 


2.88 


3.28 


2.75 


512 


9.5 


3 


4.30 


3.70 


3.70 


3.36 


3.36 


3.32 


542 


9.5 


3 


4.20 


3.36 


3.12 


3.21 


3.01 


2.88 


352 


9.5 


3 


3.10 


2.98 


3.11 


2.80 


2.95 


2.36 


702 


9.5 


6 


4.20 


3.36 


3.52 


3.07 


3.28 


2.94 


528 


9.5 


6 


4.30 


3.79 


3.61 


3.28 


3.36 


3.27 


368 


9.5 


6 


4.30 


3.52 


3.70 


3.36 


3.52 


3.16 


390 


9.5 


6 


4.30 


4.00 


3.79 


3.14 


3.01 


3.36 


690 


9.5 


9 


4.20 


3.21 


3.52 


3.21 


2.85 


2.95 


376 


9.5 


9 


4.30 


3.36 


3.44 


3.21 


3.21 


3.44 


530 


11.0 


3 


4.30 


3.36 


3.70 


3.36 


3.28 


3.11 


364 


11.0 


3 


4.30 


3.79 


3.78 


3.21 


3.28 


3.44 


378 


11.0 


3 


4.30 


3.61 


3.21 


3.21 


3.21 


3.23 


498 


11.0 


6 


4.50 


3.52 


3.70 


3.14 


3.28 


3.08 


388 


11.0 


6 


4.30 


3.44 


3.52 


3.28 


3.61 


3.44 


398 


11.0 


6 


4.30 


4.11 


3.70 


3.70 


3.44 


3.44 


510 


11.0 


9 


4.40 


3.36 


3.44 


3.07 


3.21 


3.23 


366 


11.0 


9 


4.30 


3.89 


3.21 


3.29 


3.34 


3.39 


396 


12.5 


3 


4.30 


3.44 


3.44 


3.21 


3.36 


3.52 


350 


12.5 


6 


4.20 


3.70 


3.89 


3.45 


3.63 


3.81 


380 


12.5 


9 


4.20 


3.36 


3.35 


3.06 


3.28 


3.39 



* Determined by Illinois Department of Transportation 
1 in = 25.4 mm 



37 



= i.QQ + 3.63(L1 + L2) 



5.20 



1} 8.46 L2 3.52 



G - 1p g ( Pi - i.5 } 

H = Pec slab thickness, inches 
LI = load on a single or a tandem axle, kips 
L2 = axle code, 1 for single axels, 2 for tandem axles 
PI = initial serviceability index 
P2 - terminal serviceability index 
Using the Spangler corner equation, the empirical model given by 
Equation 3.1 was modified and extended to include material properties: PCC 
flexural strength (FF), modulus of elasticity (E), and modulus of foundation 
support (k). The following basic assumptions were made in this extension: 

a. The variation in pavement life (W) for different load magnitudes 
of the same level of the ratio of tensile stress/strength of the PCC slab 
is accounted for by the basic AASHO Road Test Equation 3.1, and is covered 
in the design procedure by the traffic equivalence factors, and 

b. Any change in the ratio tensile stress/strength resulting from 
changes in the values of E, k, and F will have the same effect on W as an 
equivalent change in slab thickness (calculated by Spangler's equation) will 
have on W as per Equation 3.1. 

The resulting final structural design model is given as follows: 



log W ]8 = 7.35 log(H + 1) - 0.06 + 



G 



1.624 x 10 



+ (4.22 - 0.32 P2) log 



1 + 



FF 



(H + 1) 



8.46 



")( 



0.75 



1.133 



215. 63J'\, 0.75 18.42 



H 



r 0.25 



(3.2) 



38 



where 

W,o = number of 18 kip single axle loads to reduce the serviceability 
index from PI to P2 ' 

FF = flexural strength of the concrete slab (28-day cure, 3rd point 
loading), psi 

H = PCC slab thickness, inches 

J = load transfer coefficient 

Z = E/k 

E = PCC slab modulus of elasticity, pci 

k = modulus of foundation support, pci 
Detailed derivation of Equation 3.2 is given in Reference 61. 

NCHRP Report 128 (Ref. 62) and the AASHO Interim Guides (Ref. 61) sug- 
gest that other terms should be added to the equation to permit variation of 
subbase quality (Q), and also regional factor (R) for climate. The re- 
gional factor should be included to account for differences in frost pene- 
tration, rainfall, daily temperature variation, and other climatic factors. 
An analysis was conducted using 16 years data from the 25 sections to 
determine the ability of Equation 3.2 to predict long term performance. 
The number of 18-kip ESAL accumulated to each year (i.e., 1962, 1968, 1969, 
1971, 1972, 1974) for each project was calculated based on mixed traffic 
data measured on the site and computed load equivalency factors. This 
value was plotted versus the number of predicted 18-kip ESAL that should 
have passed over the section to cause a loss of serviceability index from 
4.5 to the value shown in Table 3.1 according to Equation 3.2. The plot 
is shown in Figure 3.1. 

The standard error of the estimate (standard deviation of residuals) 
is 0.31 for log W. The plot clearly shows that Equation 3.2 predicts 



39 



10 



6 



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CO 
LlI 



co 10 

0) 



O) 



10' 






r — 


T 


1 — 


T 


i — r 


1 1 1 




1 

A 


1 


i — 


1 1 I" 


TT7 


— 


AASHTO 


EQ. 












A 
A A 








/ - 


— 














AA 


A A 








— 


— - 














AA 
A 


A 








— 


— 














D 
D 


B |l 

D 


a 

D 


A 

d y 








- 


— 














a 


D / 








- 














k 


/%% 


















• 

• 


• 


••X 


• • 




















t 




V 


• 
• 








— 


— 














§. 










— 


— 












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• 










- 


— 


• 




D 


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8? 


• 










- 


— 


























— 


1 


t 




• 
D 












Code 


i 




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— ~ 




1 


| 


1 

| 


| 


1 1 


1 

1 1 1 




1 


1 


• 

D 
A 


9.5" 

II" 

12.5" 

1 1 1 


1 1 



10' 



10' 



10 



8 



Actual 18-kip ESAL 



Figure 3.1. Predicted Versus Actual Equivalent 18-kip Single Axle Load 

Applications using AASHO Equation 3.2 and 16 Years Data From 
AASHO Road Test Site (1958-1974). 



40 



poorly for thick slabs (H >_ 11 inches). For example, the actual 18- kip 
(80 kN) ESAL for a 12-1/2 inch (318 mm) inch slab is about 18 million, and 
the computed value according to Equation 3.2 is over 50 million. This 
standard error is considerably larger than the error based on only the 
results from the two year Road Test which was 0.22. Additional information 
is presented later to show that the equation has an incorrect "form" for 
predicting the performance of thicker slabs. 

3.3 DEVELOPMENT OF NEW EQUATIONS 

3.3.1 Modified AASHO Equation . An approach similar to that used to 
develop the original AASHO Road Test Equation 3.2 was used to develop a 
modified equation. Data from the 25 AASHO sections were used in the regression 
analyses to determine the functions 3 and p. This equation which is called 
the "modified AASHO" equation is as follows: 

log W 1R = 1.27 61 log(H + 1) + 5.9802 + ^-rjj (3.3) 

10 0.00342(H + ])^'* ,q 

The standard error of this equation is 0.158. Although Equation 3.3 has 

a much smaller standard error than Equation 3.2, the form of the equation 

does not fit the performance data for thicker slabs as is illustrated in 

Figure 3.2. This plot shows the serviceability index versus accumulated 

18-kip (80 kN) ESAL for all data available from the 11 inch (279 mm) PCC 

slabs. The data shows that although there is a loss of serviceability, the 

serviceability index levels off, but the new AASHT0 Equation 3.3 continues 

to show a rapid loss. A typical plot for a single 9.5 inch (241 mm) PCC 

slab is shown in Figure 3.3. The modified AASHO Equation 3.3 cannot fit 

this type of performance curve because of its mathematical form. Therefore, 



41 




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42 



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43 



a new approach based on a different mathematical function is needed to 
develop an equation that "fits" the form of the performance curve. 

3.3.2 New Approach . The following mathematical form better fits the 
actual measured performance curves for the 25 AASHO sections 



3 ^ . „i ln 6 



W 18 = [p lntf- - 1) + 3]10 u (3,4) 



where W-jo = total equivalent 18-kip single axle loads to reduce the 

serviceability index from PI to P2. 



3 = -50.08826 - 3.77485H + 30.64386 / H 
p = -6.69703 + 0.13879H 2 

y = P2 + — - PI 

(e)- 3/p + 1 

P2 = terminal serviceability index 
PI = initial serviceability index. 
H = PCC slab thickness, inches 

The standard error of the estimate of log W-jo 1S 0.22. The adequacy 
of Equation 3.4 to predict the performance of the 25 sections over a 16 
year period is shown in Figure 3.4 which can be directly compared with 
Figure 3.1. The data plotted in Figure 3.4 is given in Table 3.2. The 
new equation has a much smaller standard error (0.22 versus 0.31) and 
the difference can be seen visually in the figures. Individual plots 
for the 8, 9.5, 11, and 12.5 inch (203, 241, 274, 318 mm) slabs are given 
in Figures 3.5-3.8. The new equation 3.4 can be seen to fit the performance 
data much better than the original AASHO Equation 3.1. 

This expression permits only an evaluation of the effect of the PCC 
slab thickness, terminal and initial serviceability index on equivalent 
18-kip load applications. Therefore, the equation was extended so that 



44 



< 
in 

LU 



I 

00 



o 

T3 

a. 




Actual 18-kip ESAL 



Figure 3.4. Accuracy of New Performance Equation (3.5) 



45 



Table 3.2. Performance Data of 25 Sections Computed Using the New 
Equation 3.5. 



Section 


Slab 


SI 


1974 Accumulated 


Computed 


No. 


Thickness 


1974 


Actual 18-kip ESAL* 


18-kip ESAL** 


672 


8.0 


1.69 


11,641,421 


15,976,948 


658 


8.0 


1.69 


11,641,421 


15,976,948 


652 


8.0 


1.69 


11,641,421 


15,976,948 


552 


9.5 


2.87 


14,091,699 


14,358,886 


676 


9.5 


2.75 


11,641,421 


15,611,260 


512 


9.5 


3.32 


14,091,699 


10,431,283 


542 


9.5 


2.88 


14,091,699 


14,260,283 


352 


9.5 


2.96 


18,643,645 


13,496,459 


702 


9.5 


2.94 


11,641,421 


13,683,436 


528 


9.5 


3.27 


14,091,699 


10,833,032 


368 


9.5 


3.16 


18,643,645 


11,737,991 


390 


9.5 


3.36 


18,643,645 


10,113,170 


690 


9.5 


2.95 


11,641,421 


13,589,634 


376 


9.5 


3.44 


18,643,645 


9,483,560 


530 


11.0 


3.11 


14,091,699 


20,256,096 


364 


11.0 


3.44 


18,643,645 


15,124,949 


378 


11.0 


3.23 


18,643,645 


18,277,072 


498 


11.0 


3.08 


14,091,699 


20,778,232 


388 


11.0 


3.44 


18,643,645 


15,124,949 


398 


11.0 


3.44 


18,643,645 


15,124,949 


510 


11.0 


3.23 


14,091,699 


18,277,072 


366 


11.0 


3.39 


18,643,645 


15,848,616 


396 


12.5 


3.52 


18,643,645 


20,347,888 


350 


12.5 


3.81 


18,643,645 


14,278,986 


380 


12.5 


3.39 


18,643,645 


23,286,432 



* Determined from traffic data in outside traffic lane from 1958 to 1974 
** Computed using new performance equation 3.4 based on loss of 



serviceability from 4.50 to 1974 SI as given 
1 in. = 2.54 mm 



46 




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48 



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Slab = 1 






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49 




to 




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4- 

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50 



other variables could be included such as the k-value and PCC modulus of 
rupture. The Equation 3.4 was extended using the Westergaard edge stress 
equation similar to the way in which the Spangler corner equation was incor- 
porated into the original AASHO Road Test equation as described in the AASHTO 
Interim Guides (Ref. 61). The following equation was obtained: 



log 10 W 18 = log 10 Wj 8 +(3. 892-0. 706P2)log 



8.789H 



0.75 



-1+0.359 



,F28, 4 1p g [ K 

l 690 j m 0.25, n [ - /lriN 0.75x 

4 log[-^ ^40H ]_ ]+0-359 



(3.5) 



in which: 



M = /l .6a 2 + H 2 - 0.675H 
a = radius of applied edge load, inches 
F28 = modulus of rupture used in design (28 day, 3rd point load adjusted 
for variability) = FF - C(y^j)FF 
FF = mean modulus of rupture at 28 days, 3rd point load, psi 
Fcv = coefficient of variation of modulus of rupture, percent 
C = 1.03, a constant representing a confidence level of 85 percent 
Z = E/k 

E = modulus of elasticity of PCC, psi 
k = modulus of foundation support on top of subbase, pci 
The Westergaard edge load equation was used because fatigue analysis and 
field observations of distress as discussed in Chapter 4 show that the slab 
edge is the critical point. 

Based upon these results, Equation 3.5 is the best performance equation 
derived and is believed to be adequate for design purposes. Its standard 
error is 0.22 based upon 150 data points from the 25 AASHO sections which is 
the same as standard error of the original AASHO Equation 3.1. 



51 



3.4 CLIMATIC FACTOR 

Pavement data were collected from 12 plain jointed concrete projects 
located in eight states. These projects ranged in age from 6 to 34 years and 
total 18-kip (80kN) ESAL in the heaviest traveled lane up to 39 million. 

Analyses were conducted to determine if the pavements from regions 
different from Illinois showed different performance. Four general climatic 
regions were chosen based upon precipitation and potential evapotranspira- 
tion, and frost heave and freeze thaw damage as defined in Table 2.3. A 
"climatic factor" is defined as follows: 

qp _ 18(computed) ,~ fi x 

W 18(actual) 

where 

CF = climatic factor 

W-.Q = total computed number of equivalent 18,000 pound (80 kN) 
/ . ,x single axle load applications to reduce serviceability index 
^ p ' from an initial value to a terminal value determined from 
performance Equation 3.5. 

W, 8 = total accumulated number of equivalent 18,000 pound (80 kN) 
/ . , \ single axle load applications to pass over pavement, deter- 
lacLuai; mined from traffic data. 

If CF equals 1.0 then there is no significant climatic effect for a given 
pavement. The CF could theoretically range from less than 0.1 to more than 
1.0. A computation of CF for each of the pavements is given in Table 3.3. 
There is considerable scatter in each climatic region, but the results indi- 
cate that pavements perform somewhat different in the general climatic 
regions. 



52 



Table 3.3. Data Used to Compute Factors 



Project 
No. 


Climatic 
Region 


H 
(in.) 


FF 
(psi) 


k* 
(pci) 


SI 


Calc.*** 

w 18 (io 6 ) 


Act.**** 

w 18 (io 6 ) 


CF 


1-3** 


WF 


8 


690 


115 


1.7 


15.44 


11.16 


1.38 


4-6 


WF 


9.5 


690 


115 


2.8 


15.07 


11.32 


1.33 


7-10 


WF 


9.5 


690 


115 


3.1 


12.25 


13.75 


0.89 


11-14 


WF 


9.5 


690 


115 


3.2 


11.40 


17.82 


0.64 


15-17 


WF 


11 


690 


115 


3.1 


20.42 


14.09 


1.45 


18-22 


WF 


11 


690 


115 


3.4 


15.70 


18.94 


0.83 


23-25 


WF 


12.5 


690 


115 


3.6 


18.62 


19.47 


0.96 


32 


WF 


10 


750 


197 


3.0 


22.83 


35.93 


0.64 


36 


WF 


10 


600 


115 


2.5 


20.87 


18.73 


1.11 


37 


WF 


9 


650 


400 


3.9 


6.11 


6.53 


0.94 


26 


W 


9 


700 


192 


3.9 


6.08 


5.42 


1.12 


33 


W 


8 


707 


200 


3.4 


12.16 


21.74 


0.56 


30 


DF 


9 


675 


267 


4.0 


5.12 


5.12 


1.00 


34 


DF 


8 


707 


353 


3.4 


8.10 


6.45 


1.25 


35 


DF 


8 


707 


367 


3.6 


6.97 


8.83 


0.79 


27 


D 


8 


675 


450 


3.0 


10.29 


14.57 


0.71 


28 


D 


8 


675 


500 


3.0 


11.58 


39.65 


0.29 


31 


D 


9 


725 


200 


3.2 


12.57 


30.23 


0.42 



* Mean of lowest 9 months on top of subbase. 
** Pavements from the Road Test site that had similar slab thickness 
and traffic were combined. 
*** Computed from Eq. 3.5. 
**** Computed from traffic data. 

1 in = 25.4 mm 



53 



Region CF Range Mean C F 

Wet/Freeze 0.64-1.45 1.02 

Dry/Freeze 0.79-1.26 1.02 

Wet/Non-freeze 0.56-1.12 0.84 
Dry/Non- freeze 0.29-0.71 0.47 
These data indicate, for example, that similar plain jointed concrete pave- 
ments constructed in dry/non-freeze climates would last about twice as long 
as those in wet/freeze climates. Additional data are needed to further 
verify these results however. 

The following values are selected for use in design. 

Wet- freeze (WF) 1.0 
Wet/non-freeze (W) 0.9 
Dry-freeze (DF) 1.0 
Dry/Non-freeze (D) 0.6 
A plot showing predicted 18-kip (80 kN) ESAL (by Eq. 3.5) vs. the 
computed 18-kip (80 kN) ESAL (or actual) for all pavements is given in 
Figure 3.9. The total 18-kip (80 kN) ESAL data is computed based on the 
heaviest traveled lane from the time the pavement was opened to traffic 
unitl the data of the survey when the SI was determined. The computed 
18-kip (80 kN) ESAL were adjusted by the recommended climatic factors. 
The standard error of prediction of Equation 3.5 is 0.175 which is less than 
the 0.222 for the AASH0 equation. 

3.5 TERMINAL SERVICEABILITY INDEX FOR ZERO-MAINTENANCE DESIGN 

A terminal SI must be selected for zero-maintenance design. The 
value selected has a significant effect on (1) construction cost of the 
pavement, (2) probability of structural distress occurring that requires 



54 



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maintenance, (3) user costs due to rough pavements. The higher the terminal 
SI the more substantial the structure must be and therefore the hi ier the 
construction cost of the pavement. However, the higher the terminal SI 
the lower the probability of structural distress and the lower the user costs 
due to rough pavements as shown in Figure 3.11. 

The loss of serviceability of a jointed concrete pavement is caused by 
several distress types. Previous analysis (Ref. 2) indicated that linear 
cracking, faulting at joints and cracks, spalling of joints and cracks, and 
differential settlement of the slabs (causing roughness) are the major causes 
of loss of serviceability. As the serviceability index drops from its initial 
value after construction, the probability of the pavement requiring mainte- 
nance increases. A plot of serviceability index versus the percentage of 

projects showing zero-maintenance performance is shown in Figure 3.10 for 

2 2 

all 37 projects (Pavements having less than 3 ft patching/1000 ft were 

considered to be relatively maintenance-free). Pavements having servicea- 
bility indices of 3.6 or higher all showed maintenance-free performance, 
while pavements having serviceability indices of less than 2.7 all received 
maintenance. The 50th percentile for receiving maintenance is a servicea- 
bility index of 3.0. 

3.6 SENSITIVITY OF NEW PERFORMANCE EQUATION 

A sensitivity analysis of Equation 3.5 is presented to illustrate the 
effect of the several parameters. The five design parameters included are 
slab thickness, modulus of rupture of PCC, modulus of subgrade reaction, 
climatic region, and total 18-kip (80 kN) ESAL application in design lane. 
The relative effects of these variables are shown in Figures 3.12 to 3.14. 
All of these curves are developed for a terminal serviceability index of 



56 



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2.5 



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Design Example 

Speed Limit = 55mph 
Discount Rate = 6% 

ADT 75,000 to 150,000 
In 20 Years 



ADT 50,000 to 100,000' 
In 20 Years 




ADT 30,000 to 
60,000 In 20 Years 



1.5 2.0 2.5 3.0 3.5 4.0 4.5 
Terminal Serviceability Index 



Figure 3.11. Total Extra User Costs vs. Terminal Serviceability 
for Different Traffic Conditions (from Ref. 2). 



58 



100 
80 

60 



— F = 700 psi 
- P f =3.0 



40- 



<0 

o 


20 


— 


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if) 






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10 




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00 


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7 8 9 10 II 12 13 

Slab Thickness — inches 



14 



Figure 3.12. Sensitivity of New Performance Equation (Eq. 3.5) 



59 







F 


1 i i i i 
= 500 psi 


I 






k 


= 50 pci 








p t 


= 3.0 




40 








^ ^^^^ — 


30 






Dry Non- Freeze -\ ^ — ' 




■> 20 






Wet Non- Freeze ^^^ ^^* 




o 










X 










18-kip ESAL 
CD 00 o 










— 












- Wet Freeze 
Dry Freeze 


4 














— 


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j 




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. 


j9in. 10, 


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Slab Thickness- in, 



13 



14 



Figure 3.13. Sensitivity of New Performance Eq. 3.5 over Climatic Regions, 



60 



100 
80 

60 
40 



— k =50 pci 
" P t = 3.0 



WF Climate 



O 



< 
co 

UJ 



00 



20 




8 



9 10 II 

Slab Thickness — inches 



13 



14 



Figure 3.14. Sensitivity of New Performance tquation (tq. 3.5) 



61 



3.0. These plots show that each of the five parameters have a significant 
effect on pavement life as measured by the 18-kip (80 kN) ESAL. An example 
of the effect of each parameter is shown in Table 3.4 where the change in 
18-kip (80 kN) ESAL is noted due to a change in each of the four other 
parameters. A significant change in each parameter, produces a significant 
change in the 18-kip (80 kN) ESAL applications. Since a typical lane 
loading for a heavily trafficked pavement is one million 18-kip (80 kN) 
ESAL/year, these could be considered as years of pavement life. 

A graphical solution of Equation 3.5 was prepared and is shown in Figure 
3.15 for a terminal serviceability index of 3.0. An example is given to 
illustrate the use of this chart in design. Assume the following parameters 
and solve for the required slab thickness: 

Climatic Region = Wet/freeze 

Total 18-kip (80 kN) ESAL (millions) = 20.0 

k-value on top of subbase = 100 pci (271 kP/cm) 

Working stress of PCC (F28) = 600 psi (4137 kP) 

Required PCC slab thickness = 11 inches (279 mm) 
If the pavement was located in a dry/non-freeze climatic region the 
required slab thickness is as follows: 

Total 18-kip (80 kN) ESAL = 20.0 x 0.6 = 12.0 

Required Slab Thickness =9.5 inches (241 mm) 
The effect of terminal SI on user costs from vehicle operation and tra- 
vel time was previously analysed (Ref. 2). Results showed that the total 
user costs over a 20 year life span increases dramatically as the terminal 
serviceability at the 20th year is decreased from 3.5 to 2.0 as shown in 
Figure 3.11. The rate of increase is ^ery high for a terminal serviceability 
less than 3.0. 



62 



Table 3.4. Example Sensitivity Analysis of New Performance 
Design Equation 3.5. 



Design 
Parameter 



Modulus of Rupture 
psi 



Change in 
Parameter * 



Slab Thickness, ins 
Climatic Region 



600 to 800 



k-value of Foundation, 100 to 500 
pci 



9 to 11 
WF to D 



Change in 
18-kip ESAL (million) 



9.2 to 15.5 

9.2 to 14.5 

9.2 to 20.0 
9.2 to 15.3 



*Values are computed assuming a standard section of 9 in. 
slab, k-value of 100 pci, modulus of rupture of 600 psi, 
located in a wet-freeze region. 



63 



15 



% 12 

Q> 

C 



jQ 

in 



10 



8 



Zero Maintenance Design Chart for 
Plain Jointed Concrete Pavement 



Min. Allowable 
Thickness 




40 



c 
o 



= 30 



< 



I 

go 

"5 
.o 



20 



10 




Terminal Serviceability = 3.0 



Figure 3.15. Graphical Solution of Equation 3.5 to be Used for Design, 



64 



Based upon these results, and the desirability of keeping construction 
costs as low as possible, a minimum acceptable design terminal SI is in 
the range of 3.0 to 3.5. A minimum value of 3.0 is selected since it pro- 
vides a reasonable assurance that no or only minimal maintenance will be 
required before the pavement reaches this value. Also, user delay costs 
have not increased dramatically at a SI of 3.0. Much of the loss in SI 
from its initial value to 3.0 to 3.5 is believed to be from differential 
settlement or possibly heave of the foundation, hence very little surface 
deterioration should have taken place. 



65 



CHAPTER 4 
STRUCTURAL DESIGN BASED ON PCC FATIGUE 

This chapter discusses the development of a fatigue analysis proce- 
dure for plain jointed concrete slabs. The purpose of the fatigue analysis 
is to prevent cracking of the slabs which is one of the most serious types 
of distress requiring maintenance. 

4.1 FINITE ELEMENT MODEL 

An evaluation was conducted to determine the best jointed concrete 
pavement structural analysis program. Several programs were evaluated in- 
cluding the following: 

a. Finite Element Program for Concrete Slabs Using Winkler Founda- 
tion (Ref. 66). 

b. Finite Element Program for Concrete Slabs Using Elastic Solid 
Foundation (Ref. 67). 

c. Concrete Airfield Pavement Design Program (PCA) (Ref. 68). 

d. Discrete-Element Method of Analysis for Orthogonal Slab and 
Grid Bridge Flow Systems Program (Slab 49) (Ref. 4). 

e. Finite Element Analysis of Concrete Airfield Pavements Program 
(Ref. 5). 

f. Finite Element Program for Pavement Analysis (Ref. 7). 

Results from the evaluation showed that the finite element slab analy- 
sis program developed by Y. H. Huang and S. T. Wang (Ref. 66) has the most 
desirable capabilities for analysis of plain jointed concrete pavement 
slabs. This program has many important capabilities including the determination 



66 



of stresses and deflections in concrete pavement slabs with full or par- 
tial subgrade contact, variable load transfer of transverse and longitu- 
dinal joints, and the effect of thermal gradients on curling stress both 
independently and in combination with traffic load. The method is based 
on the theory of minimum potential energy by dividing the slab into small 
elements interconnected only at a finite number of nodal points. Other 
major advantages of the finite element method are that computed stresses 
agree well with experimental results, elements of varying sizes 
can be easily incorporated in the analysis, and no special treatment is 
needed at a free edge. 

4.1.1 Description of Finite Element Method . This finite-element 
method is based on the classical theory of thin plates which assumes that: 

(a) The plane before bending remains plane after bending. 

(b) The slab is homogeneous, isotropic, and elastic. 

(c) The subgrade acts as a Winkler foundation, i.e., the reactive 
pressure between subgrade and slab at any given point is propor- 
tional to the deflection at the point (dense liquid approximation 
of the subgrade). 

A set of simultaneous equations are obtained for solving the unknown 
nodal displacements of every element in the PCC slab and a force-displacement 
relationship for all nodes in the pavement model is developed as 

k[A]{6'} = {F} [k]{6} (4.1) 

where k is the modulus of subgrade reaction, [A] is the diagonal matrix 
representing the area over which subgrade reaction is distributed, {<$' } 
is the subgrade displacement, {F} is a vector containing all the forces 
acting on the pavement model, {6} contains all the nodal displacements, and 



67 



[k] is the assembled stiffness matrix of the whole pavement system (for 
more details see Ref. 66). The first term on the left hand side of Eq. 4.1 
represents the nodal forces due to subgrade displacement caused from the 
resultant of the initial curling of the pavement slab due to a temperature 
differential between the top and the bottom, and the amount of the initial 
gaps due to pumping or plastic deformation of the subgrade. The amount of 
the initial gap as well as the deformed shape of the slabs due to the com- 
bined effect of slab weight and warping are determined first. The nodal 
moments and stresses are then computed from the nodal displacements using 
the stress matrix tabulated by Zienkiewicz (Ref. 8). Because the stresses 
at a given node computed by means of one element might be different from 
that by the neighboring elements, the stresses in all adjoining elements 
are computed and their average values obtained. 

4.1.2 Transverse and Longitudinal Joints . The finite element model 
provides an effective method for analyzing concrete slabs with doweled 
transverse joints. First by assuming the discontinuity of the two adjacent 
slabs at the joint, the equilibrium equations of each slab is developed 
separately. Assuming there is no moment transfer across the joint (since 
dowel bars do not transmit much moment from one slab to the other), then 
dowel bars will effect only those equations that give vertical forces at 
each node. Therefore, by equating the sum of two equations corresponding 
to vertical forces at every two adjacent nodes at the joint, to the external 
forces applied at that node, the number of equations is reduced. However, 
at every two adjacent nodes at the joint, one equation (the efficiency e- 
quation of load transfer) is added to the set of the equilibrium equations 
which result in the total number of equations remaining unchanged. 



68 



4.1.3 Computer Program . The finite element model is programmed 
for an IBM 360 computer. The program can determine the slab stresses and 
deflections due to thermal curling (stresses caused only by weight of 
slabs), applied traffic loads, or the combination of the applied loads 
and thermal curling of the slab. The program can handle one slab, two 
slabs connected by a transverse joint, or four slabs connected by a longi- 
tudinal and a transverse joint. The efficiency of load transfer at each 
joint can be specified as the ratio of the deflection of the unloaded slab 
over the deflection of the loaded slab at the transverse joint, as a 
percent: 

r-.cc- • unloaded 1on , n \ 

Efficiency = -r. x 100 (4.2) 

W loaded 

The tire imprints of the wheel load are converted to rectangular 
areas, and the coordinates of their sides must be input so that the pro- 
gram can distribute the wheel loads among adjacent nodes by statics. The 
program can handle any number of loads at the same time. The additional 
computer time due to these additional loads is ^ery small because Gauss 
elimination of the coefficient matrix is carried out only once regardless 
of the number of loads involved. 

Computation of the thermal curling stress of the slab requires input 
of temperature differential through the slab. 

The program can be used to investigate the effect of partial subgrade 
contact on stress distribution. The nodal numbers at which subgrade reac- 
tions resulting from loss of subgrade contact does not exist can be assigned, 
and the first term on the left side of Equation 4.1 will be automatically 
eliminated at these nodal points when forming the simultaneous equations. 



69 



4.1.4 Comparison of Measured and Computed Load Stress . A comparison 
is made between the finite-element solutions and experimental measurements 
so that the validity of the method as applied to actual pavements can be 
tested. The results of the strain measurements from AASHO Road Test (Ref. 9) 
provide excellent data for making such comparisons. Tests were conducted 
on the main traffic loops where the strain due to moving traffic was measured 
at the slab edge far from any joint. The length of slabs consisted of 15 ft. 
non-reinforced sections and 40 ft. reinforced slabs and slab thickness 
ranged from 5 to 12.5 ins. 

The finite element program requires the modulus of elasticity and the 
Poisson's ratio of concrete, the modulus of subgrade reaction, k, and the 
axle load. The measured dynamic modulus of concrete was 6.25 x 10 psi , and 
the Poisson's ratio was 0.28. The determination of the subgrade k-values is 
much more difficult because it changes appreciably with the time of the 

year. The elastic k-values on the subbase obtained by the plate bearing test 

3 
at the AASHO Road Test varied from approximately 85 to 200 lb/in. over 

all of the loops throughout the two years. Two k-values of 108 and 150 pci 
were used in a F.E. analysis conducted to verify the closeness of the program 
to the measured values at the AASHO Road Test. The first value (108 pci) is 
the mean k- value that was measured during the spring trenching program between 
April 23 and May 25, 1960. The second value (150 pci) is somewhat of an over- 
all average from the loops as indicated from Figure 3-8, Reference 9. 

The single and tandem axle loading configurations are shown in Figures 
4.1 and 4.2, respectively. The load configuration shown in Figure 4.3 was 
also used and was found to give the same stresses as the configuration in 
Figure 4.1 . 

The stress comparison for single axles is shown in Figure 4.4 and for 



70 




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74 



tandem axles is shown in Figure 4.5. The distance from edge of the slab to the 
center of the wheel load was 20 in. (508 mm) in the F.E. analysis, which is 
similar to the 17-22 in. (432-559 mn) measured for the actual loadings. Com- 
pressive strain at the top of the slab was measured in the longitudinal 
direction, 1 in. (25 mm) from the edge of the slab. The strain measurements 
were correlated with axle load, PCC slab thickness, and temperature difference 
(standard differential) and regression equations were developed (see Eqs. 
78 and 79 in Ref. 9). The theory of elasticity was then used to convert 
the strain equations into stress equations (Ref. 10) as follows: 
Single Axles: 

139. 2L- 



a. 



s 1Q 0.0031%1.278\ 
Tandem Axles 



J (4.3) 



,o°- 0031t (h 



25 ' 87L 1 (4.4) 

°t n .0.0035T/,0.8523 < | 



10 0.0035T^. 






where 



a = predicted stress at edge of slab for single axle load located 
s 17-22 in. (432-559mn) from edge, psi 

a. = predicted stress at edge of slab for tandem axle load located 
17-22 in. (432-559mn) from edge, psi 

L-, = axle load of truck (a single axle or a tandem axle set), kips 

H = PCC slab thickness, inches 

T = the temperature at a point 1/4 inches (6 mm) below the top surface 
of the 6.5 in. (165 mm) slab minus the temperature at a point 1/2 
in. (113 m^n above the bottom surface, determined at the time the 
strain was measured (called standard differential). 

The axle loads used for these plots were 18-kip single and 36-kip tandem. 



75 



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76 



The results show good correlation between the stresses computed with the 
finite element program and the stresses computed with the AASHO equations 
for both single and tandem axles. Thus, the finite element program can be 
used with confidence to compute stresses caused by axle loads. 

4.1.5 Comparison of Computed and Measured Thermal Curl Stress . Results 
from the tests conducted by Teller and Southerland at Arlington, VA (Refs. 12 
and 29) provide "measured" thermal curling stress data (actually computed 
from measured strains) for plain jointed concrete pavement that can be used 
for comparison with computed finite element stress. Slabs were 20 ft. (6.1 m ) 
in length and 10 ft. (3.0 m) wide with one exception, where a 10 ft. (3.0 m ) 
long slab was obtained after a transverse crack occurred. Longitudinal edge 
stresses were computed from measured strains during periods of maximum thermal 
gradients for several days in 1934, A summary of data obtained for 9 and 
9-6-9 in. (229- 152- 229mn) slabs and the computed FE edge stress are given in 
Table 4.1. Also some data for 15 ft. (4.6 m) slabs as later reported by 
Southerland (Ref. 29) is given. These data are plotted in Figure 4.6. The 
results generally agree, however, computed stress using the finite element 
program is slightly higher than measured stress for the limited data availa- 
ble. 

The computation of curling stresses using the finite element (FE) pro- 
gram provides a much more realistic analysis than the Westergaard/Bradbury 
analysis (Refs. 17 and 18). The FE program allows the slab to curl in a 
weightless condition, and then the restraining weight of the slab is added. 
Hence, the slab is restrained by its weight. The Bradbury model assumes full 
restraint of the slab which should give higher stresses. A comparison of 
stress computed by each method over a range of slab thickness and foundation 
modulus values is given in Table 4.2. The thermal curling stress computed 



77 



Table 4.1. Measured and Computed Curling Stress at Edge of Slab.* 



Curling Stress at Edge-psi 



Thermal 

Gradient 

(°F/in) 


Slab 

Length 

(ft) 


Measured 
8 in. Slab 


Measured 
9 in. Slab 


Measured 
9-6-9 in. Slab 


Finite 
Element 
Computed 
Stress (psi) 


3.0 


20 


— 


— 


316 


307 


2.3 


20 


— 


— 


218 


235 


3.2 


20 


— 


— 


291 


327 


2.7 


20 


— 


191 


245 


276 


3.3 


20 


— 


298 


380 


338 


3.1 


20 


— 


306 


380 


317 


3.6 


20 


— 


302 


361 


368 


3.5 


20 


— 


329 


409 


358 


2.8 


20 


— 


213 


282 


286 


2.0-3.0 


10 


— 


— 


19-68 


48-72 


2.0-3.0 


15 


115-120 








136-204 



*From Refs. 12 and 29. 
E = 5 x 10 6 psi, y = 0.15, k = 200 pci 



78 





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79 



Table 4.2. Comparison of Thermal Edge Curling Stresses Computed 
Using Bradbury (Ref. 18) and Finite Element Models 
(Ref.66). 



Slab 




Fou 


ndation 


Mod 


ulus 


(k) ■ 


■ PC 


i 


Thick-in 


Model 


50 






200 






500 


8 


Finite Element 
Bradbury 


134*p 
144 


si 




204 
246 






247 
294 


10 


Finite Element 
Bradbury 


105 

98 






178 
255 






204 
330 


14 


Finite Element 
Bradbury 


66 
100 






129 

210 






145 

341 



Parameters Used: Slab Length = 15 ft. 

Thermal Grad = 3° F/in. fi 
Mod. of Elast. = 5 x 10 psi 
Coef. of Exp. = 5 x lQ- 6 /°F 



80 



from the Bradbury model is on the average 43 percent greater than the stress 
computed using the finite element model. This result is significant in that 
most of the comparisons of measured curling stress with Bradbury or Wester- 
gaard computed curling stress showed that the measured stress was less than 
the computed stress (Ref. 29). Hence, the finite element procedure is closer 
to measured curling stress than the Bradbury procedure. The Bradbury proce- 
dure gives much higher stresses for thicker slabs than the finite element 
model as shown in Table 4.2. 

There are at least two reasons why the computed curling stress is less 
than measured stress. The effect of a moisture gradient through the slab 
(dryer on top than on bottom) causes stresses of opposite sign to thermal 
curl stress (when top of slab is warmer than bottom), and hence reduced 
measured strain. Another reason is some settlement of the subgrade which 
would reduce curling stress. In summary, considering the hazards in measur- 
ing thermal curling strains, computed and measured values are reasonably 
close. However, the finite element computed thermal edge curling stresses 
are somewhat higher than the "measured" stresses (computed from strains). 

4.2 EFFECTS OF PAVEMENT FACTORS 

The finite element program (Ref. 66) provides a powerful tool for evaluat- 
ing the effect of several pavement factors on critical stress in the PCC slab. 
A sensitivity analysis is provided and the results are used several times 
throughout the chapter in the development of the fatigue analysis. 

The traffic loading includes an 18-kip single axle and a 36-kip tandem 
axle located at the edge of the slab at midpoint between transverse joints 
as shown in Figure 4.1 and 4.2. The critical stress for this load position 
is at the bottom of the slab edge, parallel to the edge beneath the wheel 



81 



load. This stress is used in all of the subsequent analyses and is referred 

to as edge stress. The stress is caused by traffic load (referred to as load 

stress ) , thermal gradient through the slab (referred to as curl stress ), or 

a combination of load and thermal gradient (referred to as load and curl stress ) 

The edge load position was used because this position was determined to 

be the most critical point for fatigue damage in Section 4.3. Stresses were 

computed for a complete factorial of six factors. 

(1) Slab thickness (H): 8, 10, and 14 ins. 

(2) Modulus of foundation support (k): 50, 200, and 500 pci 

(3) Thermal gradient (G): -1.5 (nighttime where bottom warmer 
than top of slab), 0, +3.0 (daytime), °F/in. 

(4) Slab length (L): 15, 20, 25, and 30 ft. 
(width was constant at 12 ft. 

(5) Erodability of support (ES): 0, 12, 36, and 60 ins. 

(longitudinal strip of width ES 
along outer slab edge). 
A summary of edge stresses for all possible combinations of these factors 
and levels for an 18-kip (80 kN) single axle load is given in Table 4.3 as 
computed with the finite element program. Some results for the 36-kip tandem 
axle load are given in Table 4.4. Results for only thermal curling stress 
(without traffic load) are shown in Table 4.5. Some of these results are 
illustrated considering a selected "standard pavement" and then varying each 
factor over a typical range and plotting the change in edge stress that results. 
The selected standard pavement is selected as follows: 
Slab Thickness (H) = 10 ins. 
Modulus of Foundation Support (k) = 200 pci 



82 



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89 



Slab Length (L) = 15 ft. 

Slab Width = 12 ft. 

Thermal Gradient through Slab (G) = 0°F/in. 

PCC Modulus of Elasticity = 5 x 10 6 psi 

_c 

PCC Thermal Coefficient of Expansion = 5 x 10" /°F 

Erodability of Support (ES) = in. 
Data are plotted showing the change in edge stress due to changes in H, k, 
E, G, and L, and their interactions. Results for edge stress caused by the 
combination of traffic load and thermal gradients are given Figures 4.7- 
4.11. Results showing the effect of qn edge stress at the slab bottom for 
only thermal gradients are shown in Figures 4.12-4. 1 5. A comparison of 
stresses caused by 1 8- kip single axle and 36-kip tandem axles is shown in 
Figure 4.16. A brief summary of the most significant results as illustrated 
in Figures 4.7-4.16 and from the data in Tables 4.3-4.5 is as follows: 

(1) As joint spacing (L) increases from 15 to 30 ft. (4.6-9.1 m), 
edge stress caused by thermal gradients change greatly. Stress increases for 
daytime gradients (i.e. top of slab warmer than bottom) and decreases for 
nighttime gradients (i.e. bottom warmer than top) as joint spacing increases 
(Figure 4.7b shows combined load and curl edge stress, Figure 4.12 shows curl 
stress only). Joint spacing does not have significant effect when only 
traffic load is applied (Figure 4.7b). 

(2) As slab thickness increases, edge stress caused by either 
traffic load, thermal gradient, or load and gradient combined, decreases sig- 
nificantly for slabs of about 20 ft. (6.1m) or less (Figure 4.8, 4.13b, 
Table 4.3). The conbined stress generally increases as slab thickness 
increases for slabs with length greater than about 20 ft. Slab thickness also 
interacts with other parameters such as the k-value and erodability. The 



90 



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300 



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w 

co 
a> 

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LlI 



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1 1 1 1 

H = 8 in. 




H= 10 




— 


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— 


1 1 1 


1 



15 20 25 30 
Slab Length, ft 



<D 



600 



500 
a. 
</T 400 

to 

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go 300- 

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ft 200 

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15 20 25 30 
Slab Length, ft 



& 



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300- 



200- 



100 



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Slab Length, ft 



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- 








ES 


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— 


| 


| 


1 


1 



15 20 25 30 
Slab Length, ft 



Figure 4.7. Effect of Slab Length on Total Edge Stress (load and curl) for 
Selected Pavement Single Axle Load. 



91 




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o. 


300 


</i 




l/> 




0) 




w 




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CO 




<D 




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8 10 12 14 
Slab Thickness, in. 




8 10 12 14 
Slab Thickness, in. 



Q. 
to 

a> 



CO 

a> 
UJ 



300- 




to 

a. 

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w 
CO 

a> 
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LU 




8 10 12 14 
Slab Thickness, in. 



8 10 12 14 
Slab Thickness, in. 



Figure 4.8. Effect of Slab Thickness on Total Edge Stress for Selected 
Pavement/Single Axle Load. 



92 



400- 



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300 


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100 




12 36 60 

Erodability (ES), in. 




12 36 60 

Erodability (ES),in. 




12 36 60 

Erodability ( ES) , in. 




100 



12 36 60 
Erodability (ES), in. 



Figure 4.9. Effect of Subbase Erosion on Total Enge Stress for Selected 
Pavement/Sincle Axle Load. 



93 



400 





300 


*n 




10 




0) 




w 




^_ 




CO 




0) 


200 


■o 




Ld 





00 




100 300 500 

Subgrade Modulus (k ), pci 




100 300 500 

Subgrade Modulus (k), pci 



to 

Q. 

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to 



500 



400 



co 300 



© 

-i — r 



oL^ 



t r 



= 3 n^£/«l: — 




-L 



100 300 500 

Subgrade Modulus! k), pci 




100 300 500 

Subgrade Modulus (k), pci 



Figure 4.10. Effect of Subgrade Modulus on Total Edge Stress for Selected 
Pavement/Single Axle Load. 



94 



o. 


400 






c/5 


300 


UJ 


200 




-1.5 1.5 3.0 

Temperature Gradient G 
(°F/in.) 




-1,5 1.5 3.0 

Temperature Gradient G 
(°F/in.) 



600- 



500- 



Q. 

«T 400 

«/> 

a> 



en 300- 



a> 

■o 200 



00 



0L 




k = 500 



1.5 1.5 3.0 

Temperature Gradient G 
(°F/in.) 

® 



600 








1 



-1.5 15 3.0 

Temperature Gradient G 
(°F/ia) 



Figure 4.11. Effect of Thermal Gradient on Total Edge Stress for Selected 
Pavement/Single Axle Load. 



95 



® 



Q. 

</T 

CD 

w 

CO 
0) 
X3 


400 

300 

200 

100 



-100 

-200 


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/ /( /g=3.0 
' X^k = 200 

/ 

~ k=50 

G = 
i i 1 i 


LJ 


k = 50 

> .k=200 

" V<\ G = -l.5 " 

1 1 1 1 




15 20 25 30 
Slab Length, ft. 



© 



400- 



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Cl 



£ 200 

w. 

co 

a> 
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Ll) 



100 

-100 
-200 



G=3.0 




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l 



&C 



ES = 60 



*"«£ 



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G = - 1.5 "" 5s 

_J I I L 



600 
500 
400 



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S 20 ° 



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H = 14 



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1 1- 



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_H=8 
^"- Ih = I0-| 

H= 14 

J I 



15 20 25 30 
Slab Length, ft. 



15 20 25 30 
Slab Length, ft. 



Figure 4.12. Effect of Slab Length on Thermal Curl Ldge Stress for Selecteu 
Pavement. 



96 



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400- 



- 300 

CL 

£ 200 

0) 



CO 

0) 

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100 

-100 
-200 



— 


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= 

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k = 50 

= -1.5 


200 


T=~500 - 




1 1 


| 


1 



8 10 12 14 
Slab Thickness, in. 



© 



400 



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If) 

Q. 



S 200 




-100 
-200 



G = 3.0 



ES=I2 




ES = 



ES = 60 



ES =60 - r r=."^cE ES=36 



ES = _ 



■ES =12 



G = -1.5 



600 
500 
400 



£ 300 



2 200 



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100 



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1 1 1 1 

-Li. 15 

- ~L = 20 

_ *^- ±=JO 

I I I L_ 

8 10 12 14 
Slab Thickness, in. 



8 10 12 14 
Slab Thickness, in. 



Figure 4.13. Effect of Slab Thickness on Thermal Curl Edge 
Stress for Selected Pavement. 



97 



® 



Q. 

(/) 
i/) 

CD 

w 

(7) 

CD 

o> 

LU 



300 



200-= 



00 







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k =200 
7"="50~~ 



G = 



k=_50_ 



-loo— F- 



k = 500 - 



k= 200 
G = 



1.5 



12 36 60 

Erodability ES, in. 



Q. 
to 



(/) 
0) 

LjJ 



400 
300 
200 
100 


-100 
-200 



_L=_3 
_L=_25 

J_ = 20 
L=I5 



G= 3.0 
G = 



1 



H h— 

__L_=J5 

L= 20 - 

T=~25 



12 36 60 

Erodability ES , in. 



© 



300 



10 

a. 


200 




100 


a> 

■o 

LlJ 






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H=8 



--===£ 



H =10 



G = 3.0 
G = 



H = 14 

1— 



H=8 \ 

»- H 



G= -1.5 

10 



12 36 60 
Erodability ES , in. 



Figure 4.14. Effect of Subbase Erosion on Thermal Curl Edge Stress for 
Selected Pavement. 



98 




100 300 500 

Subgrade Modulus (k ), pci 



® 



400- 



300 



S. 200 



to 
</> 

w 

a> 

UJ 



00 


-100 
-200 



i r 



L=30 ^ 
-L=25 ^. 



— 1=20 
_ /L = I5 



=f= 



G = 3.0 



G = 
I 



--L. = l5 



L=20 



L=,30 



: 



100 300 500 

Subgrade Modulus (k) , pci 



© 



300 



2. 200 

I 100 
(?) 



a> 
o> 
■o 

UJ 







-100 



6 = 3.0 



ES = 



4 

G = 




ES = 36 
ES = I2 

1 — 



ES = 60_ 




G=-l.5 



V 



ES=I2 



= 60 
S=~0 



100 300 500 

Subgrade Modulus(k ) , pci 



Figure 4.15. 



Effect of Subgrade Modulus of Thermal Curl Edge Stress for 
Selected Pavement. 



99 



._ 300 — 



o. 
to 

0) 



0) 

UJ 




en 200 — 



Q. 

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w 

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8 10 12 14 
Slab Thickness, in. 



300 


— 1 — 


- -i 

Sv ^SAL. 


! - 


200 




t./uT^ — - 




100 


1 


1 


1 



50 200 500 

Subgrade Modulus ( k), pci 



300 — 



Q. 

(A 
0) 



UJ 




(/) 200 — 



12 36 
Erodability (ES) , in. 



Figure 4.16. Comparison of Single and Tandem Axles (shown in Figs. 4, 
4.2) on Edge Stresses in Selected Pavement. 



ana 



100 



effect of change in stress for thick slabs for changes in k-value or 
erodability of support is less than it is for thin slabs (Figures 4.8a 
and 4.8b). 

(3) As erodability of support increases, combined load and curl 
edge stress increase with one exception (Figures 4.9 and 4.14). When a 
daytime gradient exists, there is a slight decrease in combined edge 
stress because of reduced restraint of the slab. 

(4) As the subgrade modulus of support increases, combined load 
and curl stresses decrease with one important exception (Figure 4.10). With 
a daytime thermal gradient (+3.0 °F/in.)» the stress generally increases as 
the k-value increases (Table 4.3, Figure 4.10c). 

(5) As the thermal gradient through the slab increases from a 
negative (nighttime) to positive (daytime) edge stress caused from both the 
combined load and curl and curl only increases very significantly (Figures 
4.11 - 4.15). When the gradient is positive (daytime) the edge stress at 
the slab bottom caused by only thermal gradient is tensile, but if the 
gradient is negative (nighttime) the edge stress is compressive. 

(6) Edge stress resulting from an 18-kip single axle load is 
approximately 15 percent greater than edge stress from a 36-kip tandem axle. 

4.3 CRITICAL FATIGUE LOCATION IN SLAB 

Location of the critical point at which cracking initiates in the PCC 
slab is vital to the development of a fatigue analysis with an objective of 
preventing slab cracking. The location of the critical point is approached 
using both field results and a comprehensive slab fatigue analysis. 

4.2.1 Initiation of Cracking - Field Results . A few road tests have 
been conducted where the cracking of plain jointed PCC slabs was carefully 
recorded. Results from the AASHO and Michigan road tests and also observa- 
tions made on in-service pavements are presented. 

101 



The AASHO Road Test provides data relative to the initiation of 
cracking. Transverse cracking occurred first on 61 out of 91 plain and 
reinforced concrete sections and "usually began with a crack originating at 
a point on the edge of the pavement at least 5 ft. (1.5m) from the trans- 
verse joint" (Ref. 9). Although longitudinal cracking initiated first in 
the other 32 sections, it usually only occurred in thin slabs (2.5-5.0 in., 
63-127 mm) and not on thicker slabs of 8 ins. (203 mn) or greater which are 
under consideration in this study. The location of the first crack in 31 
plain jointed concrete sections is as follows: 

Distance From Joint Number of 

To Transverse Crack, ft. Failed Section 

0-5 5 

5-10 20 

10-15 6 

An example of crack initiation and progression for an 8 in. (203 mm) plain 
slab section is shown in Figure 4.27. The initiation of most of the cracking 
at the slab edge near the midpoint of the slab is apparent. 

Results from the Michigan Test Road (Ref. 19) also show transverse 
cracking to be the dominant type occurring for slabs of 8 in. (203 mm) thick- 
ness having joint spacing ranging from 10 to 30 ft. (3 to 9 m). A dia- 
gram of the cracking of these slabs after 10 years of traffic is shown in 
Figure 4.18. Cracking in 1955 per mile is as follows: 

Joint Spacing, ft. 





30 


20 


15 


Transverse 


296 


139 


50 


Diagonal 


7 


6 





Longitudinal 


4 


20 


11 



102 





Cracking Index, 9ft. Per 1000 Square Feet 
Serviceability Index 4.0 




Cracking Index, 42ft. Per 1000 Square Feet 
Serviceability Index 3.2 




Cracking Index, 62ft. Per 1000 Square Feet 

Serviceability Index 2.9 

^---—Patched Area 




Cracking Index, 92 ft. Per 1000 Square Feet 
Serviceability Index 1.5 


in. (203 n 
r Subbase. 










n an 8.0 
) Granule 






, 


\ 


racking i 
n. (76 mm 
Road Test 




— 






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on 3.0 i 
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104 



Transverse cracking occurred much more than any other type of cracking. 

Several heavily trafficked plain jointed concrete pavements were 
examined during the field survey and the type of cracking noted and summarized 
in Table 2.6. Transverse cracking was observed in 12 pavements, corner 
cracking occurred in one, and longitudinal cracking occurred in three projects. 

In summary, results from field observations indicate that for slabs of 
normal thickness (i.e. ^L 8 ins. [203mm]), cracking usually initiates at the 
slab edge and propagates transversely across the slab. These cracks are 
usually located in the center third of the slab. Available data indicate 
that transverse cracking occurs much more often than longitudinal or corner 
cracking. Longitudinal cracking generally occurs about 24 to 42 ins. [0.6-1 . lm] 
from the slab edge (Ref. 9) and initiates at the transverse joint. Corner 
cracking occurred at the outer transverse joint corner on one project having 
skewed joints without any mechanical load transfer device (LTD) and where 
pumping of the subbase was evident. It was also observed on another project 
with skewed joints (no LTD) that had severe faulting and pumping at the 
joints. Corner cracking did not occur on any slab having dowels at the 
transverse joints. 

4.2.2 Initiation of Cracking - Fatigue Analysis . A comprehensive 
fatigue analysis was conducted using the finite element program and Miner's 
fatigue damage hypothesis to determine theoretically the critical point in 
the slab where cracking should initiate. Two general positions in the slab 
were evaluated based on the results from the field studies: 

(1) near the transverse joint where longitudinal cracking initiates 
(this is the critical fatigue location assumed by PCA in their 
design procedure, Ref. 49). 



105 



(2) at mid-slab between the transverse joints where transverse cracking 
initiates. 
These locations and the direction of critical stresses are shown in Figure 
4.19. 

A fatigue analysis was conducted by considering typical variations 
in truck axle weights, axle types, and lateral placement across the slab. 
The axle load distribution used is given in Table 4.6 which is typical for 
a major truck route. The lateral truck placement in the outer lane as measured 
by Emery (Ref. 12) was used. The lateral distance from the edge of the slab 
to the outside of the truck duals is designated D as shown in Figure 4.20. 
The lateral placement of trucks in the lane, or D, was found to be approxi- 
mately normally distributed (see Section 4.5) with a standard deviation of 
10 ins. (254 mm). Table 4.7 gives the lateral displacement percentage for 
every six inch (152 mm) increment across the slab which were computed using 
a normal distribution for five mean lateral placement positions (i.e. 
D = 12, 24, 36, 42, and 48 ins. [0.3, 0.6, 0.9, 1.1 , 1.2m]). The mean lateral 
displacement of truck traffic probably varies from 12 to 30 ins. (0.3 to 0.8 m) 
depending on shoulder conditions, existence of curb and gutter, and any 
lateral obstructions. For example, if the mean lateral placement was 
D = 24 ins. (610 mm) the percent of loads passing between 3 and 9 ins. 
from the edge of slab is 4.89. 

Stresses in the slab were computed using the finite element program 
for lateral positions of D = 0, 6, 12, 18, 24, 30, 36, 42, and 48 ins. 

A typical plot of tensile stress for a 
10 in. (254mm) slab placed on a foundation with k = 200 psi and an 18-kip 
single axle load at mid-slab is shown in Figure 4.21. The highest stress 
occurs when the load is at the slab edge (D = ins.) and decreases as the 



106 



0> 



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CD 







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107 



Table 4.6. Axle Load Distribution Used in Fatigue Analysis, 
Axle Load Group (kips) Percent 



Single Axle: 




<3 


7.52 


3-7 


16.82 


7-8 


8.01 


8-12 


16.56 


12-16 


4.92 


16-18 


2.01 


18-20 


0.87 


20-22 


0.20 


22-24 


0.08 


24-26 


0.01 


26-30 


0.01 


30-32 


0.01 


Tandem Axle: 




<6 


0.37 


6-12 


10.31 


12-18 


5.36 


18-24 


5.63 


24-30 


9.92 


30-32 


5.90 


32-34 


3.73 


34-36 


1.24 


36-40 


0.45 


41-42 


0.03 


42-44 


0.01 


44-46 


0.01 


46-48 


0.01 


48-50 


0.01 



Total 100.00 



108 



Table 4.7. Percentage of Truck Wheel Loads at Various Lateral 
Distances from Slab Edge*. 



Mean Lateral Displacement (D), ins 



Position on Slab 
(or Shoulder) , D 


j_2 


24 


36 


42 


48 


< -3 ins. 


6.68 


0.35 


0.01 


-- 


-- 


-3 to +3 


ins. 


11.73 


1.44 


0.04 


0.01 


-- 


+3 to +9 


ins. 


19.80 


4.89 


0.30 


0.04 


0.01 


9 to 15 


ins. 


23.58 


11.73 


1.44 


0.30 


0.04 


15 to 21 


ins. 


19.80 


19.80 


4.89 


1.44 


0.30 


21 to 27 


ins. 


11.73 


23.58 


11.73 


4.89 


1.44 


27 to 33 


ins. 


4.89 


19.80 


19.80 


11.73 


4.89 


33 to 39 


ins. 


1.44 


11.73 


23.58 


19.80 


11.73 


39 to 45 


ins. 


0.30 


4.89 


19.80 


23.58 


19.80 


45 to 51 


ins. 


0.04 


1.44 


11.73 


19.80 


23.58 


51 ins. 


0.01 


0.35 


6.68 


18.41 


38.21 



100.00 100.00 100.00 100.00 100.00 



*Data computed from normal distribution with standard deviation = 10 ins 
(254 mm), and mean D as indicated. 



109 



n 



Y / \ / 



8'-0" 



Truck Body 



Axle 



D 



\ / \ /, 



V ..., » -,. . . .f-,,- ! ; . "i .■■..■ fff ,....... ■ ...,..■■-... ....■■..■.'■■ .*.. .* 'i ' ■ i 

A^-.-^-.^:v- : »-:*o^.'.r..^>*-'-* Vcr> ■:■: .-••.;■. .&:■ ••-.■•■-•'• -.0:. • ; . :-■•.■<* •■•.;■'.•■ .-<£ •„-■.•;.■ ;■.•»..■_-':•■. 

'•••'• ■ ■ y ■■ - ■ ■ ■ -. - ■ ■ • . . ■ , 



PCC Travel Lane 



Paved Shoulder 




D = Distance From Slab Edge To 
Outside Of Dual Tires 



Figure 4.20 Illustration of the Mean Distance from Slab 
Edge to Outside of Dual Tires. 



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111 



load is moved inward (D = 6, 18, and 30 ins. [152, 457, 762 mm]). A 
similar plot is shown in Figure 4.22 for the axle load placed at the trans- 
verse joints. Here the transverse stresses are compressive for D = 0, 6, and 
18 ins. (0, 152, 457mm) due to contraflexure at the bottom of the slab. The 
magnitude of these stresses at the transverse joint are considerably less 
than for the edge loading position (57 psi versus 225 psi). A joint load 
transfer efficiency of 50 percent was used which is a typical value as 
defined by Eq. 42. when dowel bars are used. 

Stresses were computed over a range of slab thickness (8, 10, 14 ins. 
[203,254,356 mm]) and modulus of foundation support (50, 200, 500 pci) 
at e\/ery six inch (152 mm) across the slab both at mid-slab and at the 
transverse joint. Fatigue damage was computed at each 6 in. (152 mm) point 
across the slab using these stresses and Miner's accumulative damage hypo- 
thesis (Ref. 30). Damage at each point was computed using the following 
expression: 

p m n. . 
Damage = Z z ~4 (4.5) 

i=l j=l mj 



where 



n.. = number of applied axle load applications at i lateral position 

and of j magnitude 
Nij = number of allowable axle load applications determined from PCC 
fatigue curve (Eq. 4.8) 
m = total number of i lateral positions of axle load 
p = total number of i lateral positions of axle load 
i = counter for lateral load positions (i = 1 corresponds to D = ins., 
1 = 2 to D = 6 ins. [152 mm ] , i = 3 to D - 12 ins. [ 305 mm ], 
... i = 6 to D = 48 ins. [1219 mm ]). 



112 



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113 



j = counter for axle load groups (j = 1 corresponds to 3-kip SA 
load, j = 2 to 6 kip SA load, etc.). 
Damage at a given point is first summed over the axle load distribution con- 
sidering the load to be located at D = ins. Then the load is shifted to 
D = 6 ins. (152 mm) and fatigue damage is summed over the axle load distribu- 
tion, and so on for D = 12, 18, 24, 30, 36, 42, and 48 ins. (0.3, 0.5, 0.6, 0.8, 
0.9, 1.1, 1.2 m). The number of applied traffic loads at each load position 
are computed according to the Table 4.7 based on a total of 10 million axles 
total to pass over the slab. 

The procedure was computerized and results from the analysis for mid- 
slab are given in Figures 4.23-4.25 for 8, 10, and 14 in. (203, 254,427mm) slabs 
with a foundation support of 200 pci. These curves show that when the mean 
lateral placement (D) is less than 36 ins. (0.9 m), the critical fatigue damage 
point is at D = or the slab edge . Similar results are obtained when slab 
thickness is held constant and the foundation support is varied from 50 to 
500. Therefore these fatigue analyses show that at a crosssection at the 
midpoint, cracking should definitely initiate at the outer edge of the slab . 
This was observed to occur in the field studies documented in Section 4.2.1 . 

Results from the analysis at the transverse joint are given in Figures 
4.26-4.28 for slab thickness (H) of 8, 10, and 14 ins. (203, 254, 427 mm) placed 
on a foundation of k = 200 pci. A load transfer efficiency of percent (a 
free edge) was used to give the most critical situation. The critical 
fatigue damage location is 24 to 42 ins. (0.6-l.lm) from the corner. 

A final important result is the difference in magnitude of damage for 
the mid-slab location versus the transverse joint location. The mid-slab 
edge position has a much higher fatigue damage than the transverse joint posi- 
tion when both were subjected to the same average lane traffic . Thus, transverse 



114 





1 1 1 

H = 8 in. _j 


— 


k = 200 pci / 


/ 


I / D = 12 in/ 


y #* J 1 

D / /I 


— 


/ ?7 / 


— 


^~^^y ?&l 


-ft// 
V 


/ "-' \48 _ 




\ 
\ 

v 




\ 

\ 
\ 

\ 




1 1 



10' 
10' 

io : 
io : 

10 



10 



-10 



-2 



-10 



-3 



10 



-4 



10 



-5 



10 



-6 



10 



-7 



10 



-8 



10 



-9 



48 36 24 12 

Distance From Slab Edge, in. 



4* 

w 

(D 

o 

E 
o 

Q 

a> 

o> 

o 
U- 



Figure 4.23. Computed Fatigue Damage across Slab Due to Lateral Distribution 
of Trucks on Lane--at Midpoint between Transverse Joints. 



H = 10 in. 
k = 200 pci 



10 



10' 







O 

E 
o 

Q 

O 



48 36 24 12 

Distance From Slab Edge, in. 



Figure 4.24. Computed Fatigue Damage across Slab Due to Lateral Distribution 
of Trucks in Lane--at Midpoint between Transverse Joints. 



116 



H = 14 in. 
k = 200 pci 



10 



10' 

io : 




n^8 



~t— === ! 



10 



-9 



w 

o> 
o 

E 
o 

Q 
a> 

o 



48 36 24 12 

Distance From Slab Edge, in. 



Figure 4.25. Computed Fatigue Damage Across Slab Due to Lateral Distribution 
of Trucks in Lane--at Midpoint between Transverse Joints. 



117 



H = 8 in. 
k = 200 pci 
Jt. Eff. = Cf % 




48 36 24 12 

Distance From Slab Edge, in. 



Figure 4.2b. Computed Fatigue Damage across Slab Due to Lateral Distribution 
of Trucks in Lane— at Transverse Joint (H = slab thickness, 
k = modulus of foundation, D = distance from edge of slab 
measured toward center of slab). 



118 




48 36 24 12 

Distance From Slab Edge, in. 



Figure 4.27. Computed Fatigue Damage across Slab Due to Lateral Distribution 
of Trucks in Lane--at Transver<;p .lnint. 



119 



H =14 in. 
k = 200 pci 
Jt. Ef f. = % 



48 



/~\ 




^J 




10 



-4 



10 



-5 



c 




I0" 6 W 




o, 
Fatigue 



v-10 



48 36 24 12 

Distance From Slab Edge, in. 



Figure 4.28. Computed Fatigue Damage across Slab Due to Lateral Distribution 
of Trucks in Lane--at Transverse Joint. 



120 



cracking would theoretically be expected to occur long before longitudinal 
cracking occurred. This result was verified at the AASHO Road Test and from 
the field survey where transverse cracking occurred much more often than 
longitudinal cracking. Hence, both field and analytical results indicate 
that for normal highway loadings and slab widths the critical fatigue damage 
point is at the slab edge. 

An analysis of stresses occurring for slab edge loading and joint 
loading further illustrates the fatigue damage results. The stresses in 
Figure 4.29 were computed with the finite element program for the loadings and 
slab properties indicated. The critical stress (vertical ordinate) represents 
the maximum stress occurring anywhere in the slab for the given loading. For 
example, when D = (load at the midslab edge), the critical stress is 390 psi 
and occurs beneath the load. If the load is moved inward to D = 24 ins. (0.6 m) 
the critical stress is 190 psi which occurs beneath the wheel load. The cri- 
tical stresses occurring when the load is at the transverse joint are highly 
dependent on the joint load transfer efficiency (L). For L = percent 
no load is transferred across the joint and the joint becomes essentially a 
free edge transferring no moment or shear. A critical stress of 270 psi 
occurs when D = and is located at the top of the slab approximately 5 ft. 
(1.5 m ) from the joint. It is important to note that the maximum critical 
stress at the mid-slab edge loading (390 psi) is much greater than the 
maximum critical stress near the joint (270 psi) for any lateral distance, 
D. 

4.4 EFFECT OF JOINT SPACING ON CRACKING 

Plain jointed concrete pavements have been constructed with joint spacings 
ranging from 10 to 30 ft. (3 to 9 m) . Available field data indicate that 



121 




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122 



joint spacing has a significant effect on transverse cracking of the slabs. 
Road tests were conducted in Minnesota and Michigan where the transverse 
joint spacing varied from 15 to 30 ft. (4.6-9.1 m) and 10 to 30 ft. (3.0-9.1 m) 
respectively. Figure 4.30 shows transverse cracking versus joint spacing 
after 10-15 years of service for 8 in. (203 mm) slabs. Both road tests show a 
dramatic increase in transverse cracking with increase in joint spacing with 
a leveling out after about 25 ft. (7.6 m). The Michigan study concluded that 
". . . joint spacing of approximately 10 ft would be necessary to completely 
prevent transverse slab cracking. The rate of transverse cracking increased 
approximately in relation to the square of increased slab length over 10 ft" 
(Ref. 9). The Minnesota study concluded "contraction joints should be placed 
at intervals of 15 ft. in order to obtain the best overall performance of 
the pavement slab from the standpoint of joint movement, cracking, warping, 
faulting, and roughness" (Ref. 10). The effect of joint spacing on trans- 
verse cracking was observed in four of the states visited during the field 
survey (California, Colorado, Washington, and Utah) where in randomly spaced 
joints of 12 to 19 ft. (3-6 m), the 18-19 ft. (5.5 m) slabs transversely 
cracked much more than the shorter slabs. Washington reduced their joint 
spacing to less than 15 ft. (4.6 m) because of transverse cracking on the 
longer slabs. Colorado reported that about 1/4 slabs that are 18-19 ft. (5.5 m) 
long crack transversely, but much less cracking occurs on shorter slabs. 
Numerous projects in California show much greater amounts of transverse 
cracking in the 18-19 ft. (5.5 m) slabs than the 12-13 ft. (4 m) slab. Trans- 
verse cracking was also observed in the 18-19 ft. (5.5 m) slabs in Toronto, 
Ontario where joint spacing ranged from 12 to 19 ft (3-6 m). 

It is very important that the reasons for this occurrence are understood 
so that it may be considered in design. There are at least four factors 
which may be contributing to the increase of transverse cracking with 

increased joint spacing. 

123 




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124 



1) Concrete Drying Shrinkage . PCC contains considerably more water at 
placement than is needed to hydrate the cement. With time, the exposed PCC 
slab loses considerable water which results in shrinkage. Water content, 
type and gradation of aggregate, chemical admixtures, moisture and tempera- 
ture conditions have all been found to affect drying shrinkage (Refs. 21, 
22, 23). The shrinkage of the PCC slab is resisted by friction at the slab/ 
subbase interface and thus tensile stresses develop in the slab. Also, the 
temperature of the PCC drops during the first night due to a reduction in 
hydration rate and lower nighttime temperature resulting in additional tensile 
stress. Transverse cracks occur generally within the first day after place- 
ment ranging from 40 to 150 ft. (12 to 45 m) in slabs that are placed with- 
out any joints. However, when joints are placed in the PCC, tensile stresses 
due to drying shrinkage are reduced due to joint movement and their effect 
becomes smaller as joint spacing decreases. Many slabs ranging up to 80 ft. 
(24 m ) in length were observed during the field survey that did not have 
any transverse cracks for pavements ranging up to 25 years in age. Hence, 
it is believed that for relatively short joint spacings (i.e. £20 ft.) [6 m] 
tensile stresses caused by drying shrinkage are mostly relieved through 
joint opening, and therefore have only minor effect, if any, on transverse 
cracking. 

(2) Concrete Temperature Shrinkage . After the PCC hardens and the 
temperature drops, both during nighttime and seasonally, tensile stress 
occurs due to thermal shrinkage of the PCC and subsequent frictional re- 
sistance of the subbase. The maximum value of these stresses can be crudely 
estimated using the "subgrade drag theory" where the slab is pulled over the 
foundation and the stress required is computed. Frictional resistance tests 
have been conducted by several researchers (Refs. 24, 12, 25) which show that 



125 



the coefficient of frictional resistance is not a constant, but increases with 
increasing displacement of slab, until a maximum is reached where the slab 
slides freely. Sliding friction values ranging from less than 1.0 to over 
2.0 have been obtained depending on foundation conditions. Stresses computed 
using conventional subgrade drag theory are given in Table 4.8 for a range of 
conditions using Equation 4.6 (Ref. 18). 

Lfy c 

where S = tensile stress in PCC, psi 

f = coefficient of frictional resistance 

L = slab length, ft. 

y = unit weight of PCC, pcf 

The tensile stress increases linearly with joint spacing which probably 

contributes somewhat to the increase in cracking with increased joint spacing. 

These stresses are very small, however, for short joint spacings (i.e. <20 ft. 

[6 m ]), but are of significant magnitude for longer joint spacings. They 

are believed to be negligable for relatively short slabs. 

(3) Slab Moisture Gradients . Hatt (Ref. 16) reported in 1925 that 

slab warping was cuased when moisture differences existed between slab surfaces 

"After curing under water (for 30 days), the slab was dried by cir- 
culating air both above the slab and through the subgrade. The more 
rapid drying of the top surface warped the slab to a maximum upward 
deflection of 0.12 in. at the corners and 0.05 in. at the unbroken edge 
during a period of 40 days, after which time upward movement ceases, . . . 
After 3 months of further drying, the slab had become approximately uni- 
formly dry throughout its depth and substantially level (or flat), . . . 
In order to saturate the bottom of the slab, water was then introduced to 
the level of the top of the gravel subgrade (opengraded) . During 110 days 
of saturation the slab again warped upward 0.20 in. at the corners and 
0.066 in. at the unbroken edge. . . The surface was (then) covered with 
saturated burlap to simulate the effect of prolonged rain. . . In 30 days 
the corner ahd dropped 0.06 in., to a deflection of 0.14 in." (Ref. 16). 



126 



Table 4.8. Computed Tensile Stress in PCC Slabs Due to Temperature 
Reduction and Subbase Frictional Resistance.* 



Joint 

Spacing (ft) 



Coefficient of Frictional Resistance 



1.0 



1.5 



2.0 



10 
15 
20 
30 
50 
100 



5 psi** 

8 

10 
16 
26 
52 



12 
16 
23 
39 
78 



10 
16 
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31 
52 
104 



* PCC unit wt. = 150 pcf 
** Computed from Eq. 4.6. 



127 



The weight of the slab, resistance from the subgrade, and any resistance 

at the slab edges apply restraint to warping of the slab, and thus stresses 

occur at the top and bottom of the slab. The stresses caused by moisture 

gradients through the slab are referred to as warping stresses. A field 

study into the effects of moisture gradients was conducted by the Bureau of 

Public Roads (Refs. 11-15) in the 1930's. Conclusions reached from these 

studies are briefly summarized as follows: 

"The data indicate that the curvature caused by moisture is princi- 
pally an upward warping of the edges caused by a moisture loss from the 
upper surface of the pavement. The downward warping of the edges, re- 
sulting from a condition in which the moisture content in the upper part 
of the pavement exceeds that in the lower part, seems to be considerably 
smaller for the conditions of these tests. . . The edges of the slab 
reach their maximum position of upward warping from this cause during the 
summer and the maximum position of downward warping during the winter, 
the extent of the upward movement apparently exceeding that of the 
downward movement considerably" (Ref. 12). 

These results indicate that the warping of a pavement slab from moisture 
gradients is mostly a seasonal change occurring over a considerable time 
interval. For example, during the winter months in Phoenix, Arizona, when 
the relative humidity is very low, the PCC slabs are noticeably warped 
upward due to the severe drying of their surface. The warp is less during 
other seasons. The Arlington tests (Ref. 12) also showed that as the sea- 
sonal warping occurs, the slab settles somewhat into the subgrade, thus re- 
ducing the restraint due to slab weight. Any creep of the concrete would 
also tend to reduce the stress from moisture warping. 

Because of the many difficulties involved (i.e. inability to measure 
moisture contents in slabs, settlement, creep, etc.) attempts have not yet 
been successful to compute or measure strains due to moisture gradients, and 
their relative magnitudes are unknown. However, the following conclusions 
appear justified based upon the available information: 



128 



(a) The top of the slab is usually dryer than the bottom through most 
of the year, causing some compressive stresses at the bottom of 
the slab. 

(b) These stresses are greater during the warm weather portion of the 
year becuase of a drier slab surface. 

(c) Moisture warping stresses at the slab bottom are generally of oppo- 
site sign than load stresses, and hence tend to reduce the combined 
stresses occurring at the slab edge or interior. 

(d) There is not enough information presently available to consider 
moisture gradient warping stress in design, however, as joint 
spacing increases, the warping stress will also increase similar 
to thermal curling stress as subsequently described. 

(4) Slab Thermal Gradients . A difference in temperature between the top 
and bottom of slabs has been shown both experimentally and analytically to 
cause significant stress in the slabs (Refs. 2, 7, 8, 17, 18, 19). The most 
recent and accurate analytical method to compute these stresses is the 
finite element model (Ref. 66). These stresses are referred to as curling 
stresses. The most critical condition occurs when the top is warmer than 
the bottom (daytime), which results in tensile stresses at the bottom of the 
slab that are additive to load stresses. The daytime stresses have been 
shown to be much larger (about 3 times) than the nighttime stresses. Teller 
and Southerland (Ref. 12) concluded the following based upon their exten- 
sive studies into thermal curling stresses in the 1930's: 

"The data that have just been presented clearly indicate that the 
stresses arising from restrained temperature warping equal in importance 
those caused by the heaviest wheel loads. The stresses from this cause 
are actually large enough to cause failure in concrete of low flexural 
strength, and since the direction of the stresses is such that they 
become added to the critical stresses caused by wheel loads, there is 
little doubt but that warping stress is primarily responsible for much 
of the cracking in concrete pavements" (Ref. 12). 

The effects of joint spacing on thermal gradient curl edge stress as 

computed with the finite element program was shown in Section 4.2, and is 

illustrated for a specific slab in Figure 4.31 with slab length varying from 



129 



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10 to 33 ft. ( 3 to 9 m ) . There is a dramatic increase in edge stress 
(which is parallel to the longitudinal slab edge) as slab length is increased 
up to about 25 ft. ( 8 m ). It then levels off as shown. This relation- 
ship is similar to Figure 4.31 for transverse cracking of slabs varying from 
10 to 30 ft. ( 3 to 9 m). While there may be a basis to conclude that 
calculated curl stress is not as large as actual curl stress (due to the oppo- 
site effect of moisture and slab settlement into subbase), the general trend 
of curl stress relates very closely to the trend of transverse cracking. 
But most importantly, the stress resulting from both combined load and curl 
for a daytime temperature gradient (top warmer than bottom) also shows the 
same trend as cracking with increasing slab length. 

These curl stresses caused by thermal gradients are of sufficient 
magnitude to cause transverse cracking even independently over a long 
time period. Loop 1, the non-trafficked loop at the AASHO Road Test, was 
surveyed after 16 years, and most of the 40 ft. ( 12 m) long slabs were 
cracked, but none of the 15 ft. ( 5 m) slabs were cracked. Possibly the 
repeated daily thermal gradient cycles over many years resulted in a 
thermal curl stress fatigue of the slabs. Other stresses such as temperature 
shrinkage and moisture gradients may also contribute however. Several 
researchers have concluded that thermal curling stress is of significant 
magnitude to contribute to transverse cracking of slabs (Refs. 12, 31, 29). 
Based upon available results it is concluded that thermal curl stresses are 
of sufficient magnitude to cause an increase in transverse cracking as joint 
spacing increases, and must be considered in any design analysis that attempts 
to prevent slab cracking. 

In summary, field experience indicated that as transverse joint spacing 
increases from 10 to 30 ft. ( 3 to 9 m ), transverse cracking increases 



131 



dramatically. The major cause of this increase is increased stress re- 
sulting from restrained thermal curling with increased joint spacing. Drying 
shrinkage and temperature shrinkage do not have a significant effect on short 
slabs because of stress relief effected by joints. The effect of moisture 
gradients is not fully known, but is believed to be usually of opposite sign 
of stresses caused by traffic load for edge and interior slab positions. 
Thus, it tends to reduce the critical combined stress effect of load and ther- 
mal curl (when top is warmer than bottom). 

4.5 DEVELOPMENT OF FATIGUE DAMAGE ANALYSIS 

A comprehensive PCC slab fatigue damage analysis was developed based 

upon the following: 

-The critical fatigue damage location in the slab is at the slab 
longitudinal edge midway between transverse joint. 

-Critical edge stresses caused by both traffic loads and thermal 
gradient curl are considered to prevent transverse cracking. 

-Both load and thermal curl stresses are computed using a finite element 
program which has been shown to provide accurate results. 

-The proportion of traffic occurring near the slab edge is used in the 
fatigue analysis. 

-Concrete strength changes with time and thus the fatigue analysis must 
be time dependent. 

-Fatigue "damage" is computed according to the Miner hypothesis. 

-A correlation between computed fatigue "damage" and measured cracking 
was determined and limiting "damage" for zero-maintenance design 
selected. 

4.5.1 PCC Fatigue . Several laboratory studies have shown that plain 

PCC beams experience fatigue failure when subjected to high repetitive 

flexural stresses (Refs. 32-41). Also, several road tests and many inser- 

vice PCC slabs have been observed to experience fatigue failure when subjected 

to many applications of heavy truck traffic (Refs. 9, 45). However, no 



132 



correlation between laboratory and field fatigue results has been 
attempted. 

The results from laboratory studies provide significant information about 
the fatigue properties of PCC applicable to pavement fatigue conditions. 

(1) The number of repeated loads that PCC can sustain in flexure be- 
fore fracture depends upon the ratio of applied flexural stress to the 
ultimate static flexural strength or modulus of rupture. 

(2) PCC does not have a fatigue limit within 20 million load applica- 
tions, hence there is no limiting repeated stress below which the life will 
be infinite (Refs. 32,33, 34). The mean fatigue strength of PCC, which is 
the strength expressed as a percentage of the static ultimate strength, is 
approximately 55 percent at 10 million applications of load (Ref. 38). 

(3) The range of loading (expressed as the ratio of flexural stress 
at minimum load divided by stress at maximum load) affects fatigue strength 
as shown in Figure 4.32. As the range increases the fatigue strength in- 
creases (Ref. 38). 

(4) Application of varying flexural stress levels gives different 
fatigue results depending on the sequence of applied loads of varied 
intensity (Refs. 34, 35). Thus, Miner's damage hypothesis, which assumes 
linear accumulation of damage, does not give exact prediction of failure 

of PCC. However, data from recent tests indicate that the inaccuracy of the 
Miner's hypothesis is not \/ery significant compared to the large variability 
in strength and fatigue life that is typical of PCC (Ref. 43). Hence, it 
was concluded that Miner's hypothesis represents the cumulative damage 
characteristics of concrete in a reasonable manner (Ref. 43). 

(5) Variability of fatigue life of PCC is very high with coefficients 
of variations ranging up to 100 percent. An example of the large variability 



133 




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of the number of applications to failure of similar specimens tested under the 
same conditions is shown in Figure 4.33. 

(6) Repeated rest periods during a fatigue test increase the fatigue 
strength, hence some recovery occurs during the rest periods (Ref. 35). 

(7) The effect of moisture conditions of PCC under flexural fatigue 
has not been fully determined. Some limited tests, indicate that high 
moisture content gives lower fatigue strength (Refs. 34, 44). However, a 
recent limited study tends to indicate that saturation affects fatigue life, 
but does not significantly change the fatigue strength (Ref. 42), however, 
the results are not conclusive. 

(8) The increase in modulus of rupture of PCC with time has significant 
effect on increasing fatigue life, but not on fatigue strength as long 

as the modulus of rupture at the specific time is used to compute the 
stress ratio (Ref. 42). This result is shown for PCC ranging in age from 
4 weeks to 3 years in Figure 4.34. The various curves for different ages 
overlap each other. 

Fatigue data were obtained for plain PCC beams from three studies 
(Refs. 33, 42, 43). A S-N plot of 140 tests from these studies is shown in 
Figure 4.35. The plot shows a large scatter of data, and the data from the 
three studies generally overlap each other. A least square regression curve 
was fit through the data as shown. 

log 1Q N = 17.61 - 17.61 (R) (4.7) 

where 

N = number of stress applications to failure of beam. 

R = ratio of repeated flexural stress to modulus of rupture 

Standard error = 1.40 (of log N). 



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This equation is a mean regression curve in that it represents a 
failure probability of 0.50 or 50 percent. Hilsdorf and Kesler (Ref. 35), 
for example, established curves for various probabilities of failure based 
upon their data, and the curve for a probability of 0.05 is plotted in 
Figure 4.36. The fatigue curve used in design by PCA is also shown in Figure 
4.36 which is much lower than the P = 0.05 curve for lower stress ratios. 

The applicability of these laboratory fatigue results from beam speci- 
mens to the fatigue of actual pavement slabs under field conditions has 
never been established. Many differences exist between laboratory and field 
conditions that probably result in different fatigue responses. Table 4.9 
has been prepared based upon a review of literature to show the probable 
differences in fatigue life of laboratory beams and of field slabs. For 
example, the effect of thermal curl which occurs daily in field slabs but 
not in laboratory beams reduces the fatigue life of slabs due to the additive 
effect of load and curl stress when the top of the slab is warmer than the 
bottom of the slab. On the other hand, the effect of age of slab and the 
corresponding increase in static ultimate strength would cause an increase 
in fatigue life of the field slab (loaded over many years) over the laboratory 
beam that is tested in a relatively short time period. The relative signifi- 
cance of each factor is indicated. Several of these factors can be reasonably 
considered in design including thermal curl, thickness variation, loss of 
support, age of PCC, and variation o^ ?CC strength. There is not adequate 
data available to directly consider the other factors, however, several 
only have minor effects and also since some have beneficial and some detri- 
mental effects they tend to cancel each other out. In summary, the com- 
plexities are so great and available information so limited that any 
laboratory curves used to estimate fatigue damage in field slabs must be 



139 




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Table 4.9. Differences in fatigue life of PCC beams (Nb) and pavement slabs 
due to certain slab conditions. 



Field Slab Condition 



Fatigue Life 



Relative 
Significance 



Variability of PCC strength 

Thermal Gradient Curl 

Thickness Variations 

Loss of Support 

Durability of PCC 

Thermal Shrinkage 

Moisture Gradient Warp 

Rate of Loading 

Rest Periods 

Age of PCC (strength gain) 

Stress Ratio 

Moisture of Slab 

Scale Effects 

Drying Shrinkage 

Load Effects (testing machine 
vs. truck traffic) 



Ns < Nb 
Ns < Nb 
Ns < Nb 
Ns < Nb 
Ns < Nb 
Ns < Nb 
Ns > Nb 
Ns > Nb 
Ns > Nb 
Ns > Nb 
Ns > Nb 
Unknown 
Unknown 
Unknown 

Unknown 



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** Moderate 



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Major 



141 



"calibrated" based on field data as presented in Section 4.6. Additional re- 
search on PCC slab fatigue is greatly needed. 

A fatigue curve must be selected for design purposes. Figure 4.36 
shows four curves plotted with the laboratory fatigue data. The mean re- 
gression curve, Eq. 4.7 essentially represents a 50 percent failure pro- 
bability curve, because approximately one-half of the beams fail before 
this curve is reached for any given stress ratio. A curve desired by Hilsdorf 
and Kesler (Ref. 35) for varying flexural stresses representing a probability 
of failure of 5 percent is shown along with the current PCA design curve. 
A curve with a fatigue limit such as the PCA curve is not selected since 
several researchers have concluded that concrete does not have a fatigue 
limit within 20 million load applications (Refs. 32, 33, 34). After consi- 
deration of a number of factors the following curve was selected for design: 

log N d = 16.61 - 17.61 (R) (4.8) 

This expression provides a safety margin of one decade of load applica- 
tion as shown in Figure 4.26, and represents a failure probability of 24 
percent. The use of a lower probability of failure curve is not felt 
justified because muc h of the variation in fatigue life is due to the inability 
to predice ultimate static strength. Since the variation of strength will 
also be considered directly in design, the fatigue curve represented by 
Eq. 4.8 is believed in reality to represent a much lower failure probability 
if variation due to concrete strength were removed. 

4.5.2 PCC Strength Increase . PCC strength data were obtained from five pro- 
jects located in as many states (Refs. 46, 47, 48, 50, 51) and also from the 
Portland Cement Association (Ref. 49). The modulus of rupture at various 
times ranging from 3 days to 17 years was obtained from tests on beams cast 



142 



during construction and then cured over time, and beams cut from slabs over 
time. The modulus of rupture was also estimated from cylinders cast during 
construction, and cores cut from the slabs over time. The cylinders and cores 
were tested for compressive strength and converted to a modulus of rupture 
using the following relationship (Ref. 52): 

Mod. of Rupture (psi) = 10[Comp. strength (psi )] ' 2 (4.9) 
Data were obtained ranging from 3 days to 17 years and many points between as 
shown in Figure 4.37. The following equation was obtained using multiple 
regression techniques. 



F A = 1.22 + 0.17 log 1Q T - 0.05(log ]0 T) 2 (4.10) 



where 



F A = ratio of the modulus of rupture at time T to the modulus of 
rupture at 28 days. 

T = time since slab construction in years. 

The modulus of rupture can be estimated at any time T using the following 

expression: 

F = F A (F 28 ) (4.11) 

where 

F = modulus of rupture at time T 
F ? n = modulus of rupture at 28 days (3rd pt. loading). 

Equations 4.10 and 4.11 can be used to estimate the ultimate static 
modulus of rupture at any time over the life of the pavement for use in the 
fatigue damage analysis. 

4.5.3 Lateral Truck Distribution . The lateral placement of trucks in the 
traffic lane is very crucial because of the high longitudinal edge stresses 
that develop when the wheel load is near the edge. The encroachment of loads 



143 



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onto the shoulder is also important in shoulder design. The lateral dis- 
tribution varies with several factors such as width of traffic lane, loca- 
tion of edge stripe, paved or unpaved shoulders, any edge restraints such as 
retaining walls, and the existence of curb and gutter. 

The lateral distribution should be measured for local conditions, and 
only general guidelines are given as to the typical range of the mean distance 
If heavy trucks travel on the average down the center lane, the mean distance 
D as shown in Figure 4.20 would be 24 ins. (.7 m ) (for a 12 ft. [3.6 m ] wide 
lane and an 8 ft. [2.4 m] wide truck). However, available evidence indicates 
that when there is a paved shoulder and no lateral obstructions, there is a 
definite tendency of trucks to shift about 3 to 12 ins. (75-305 mm) toward 
the slab edge, which gives a mean value for D of approximately 12 to 21 ins. 
(25-533 mm ). Bureau of Public Roads measurements at 15 locations in 1956 
for 12 ft. (3.6 m) concrete traffic lanes and paved shoulders on two lane 
rural highways showed an average of 11 ins. (279mm) (Ref. 53). Studies 
by Emery in 1975 (Ref. 54) showed a mean of approximately 16 to 18 ins. 
(406-457 mm) on rural four lane Interstate highways. The lateral distri- 
bution obtained from Emery (Ref 54) is approximately normal with a standard 
deviation of 10 ins. (25 mm) as shown in Figure 4.38. 

Data were collected by Taragin (Ref. 53) in 1956 for 12 ft. (3.6 m) 
concrete traffic lanes with (1) contrasting bituminous shoulders, and 
(2) grass or gravel shoulders. The average placement from the center of 
truck to the centerline paint strip was 7.1 ft. (2.1 m ) for bituminous 
shoulders and 5.9 ft. ( 1.8 m ) for grass or gravel shoulders. These 
placements correspond to the following mean lateral distances, D: 
Bituminous Shoulders = 0.9 ft. or 11 ins. (two lanes) 

Grass of Gravel Shoulders = 2.1 ft. or 25 ins. (4 lanes) 



145 





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". . . for traffic lanes of the same width, the placements on sections 
of bituminous shoulders are considerably farther from the centerline 
of the pavement than on sections with grass of gravel shoulders. The 
results of this study confirm the results of earlier studies that bitu- 
minous-paved shoulders which appear distinctly different from the traffic 
lane increase the effective pavement width at least 2 ft., regardless of 
lane width" (Ref. 53). 

Surprisingly, the lateral distribution data from the bituminous shoulders 
were not used and the general distributions for gravel shoulders was used 
to program the lateral off-set for the AASHO Road Test. However, nearly 
all heavily trafficked highways have paved shoulders, and hence the propor- 
tion of edge loads and shoulder encroachments are usually much greater than 
the programed loads at the AASHO Road Test. Results from the study by 
Emery (Ref. 54) show that approximately 10 percent of loads were within 6 in. 
(152 mm) of the slab edge on rural Interstate highways with paved shoulders. 
Several observations made by the author tend to confirm that there are a sig- 
nificant proportion of loads near the edge. Additional data are greatly 
needed to establish lateral placements of trucks for varying conditions. 
4.5.4 Thermal Gradients . The thermal gradient in the PCC slab is defined 
as follows: 

r - top bottom r« no) 

where G = thermal gradient, °F/in. 

T, = temperature at the top of the slab, °F 
T. tt = temperature at the bottom of the slab, °F 
H = PCC slab thickness, inches 

A positive gradient indicates the top of the slab is warmer than the bottom 
which normally always occurs during the daytime. A negative gradient indi- 
cates that the bottom is warmer than the top which normally occurs during the 



147 



nighttime. A positive gradient results in tensile stress at the bottom of 
the slab, and a negative gradient results in compressive stress at the 
bottom of the slab. During times when the gradient is positive (termed 
daytime) the total combined stress at the bottom of the slab edge midway 
between the joints under traffic load will be much greater than when the 
gradient is negative (termed nighttime). 

The temperature gradient varies continually throughout a 24-hour daily 
period and also varies from month to month. A mean monthly positive gra- 
dient (called mean daytime gradient) can be used in design. An illustration 
of the mean daytime and nighttime gradients for northern Illinois is shown in 
Figure 4.39 for two slab thicknesses. A summary of mean monthly thermal 
gradients for the four climatic regions is given in Table 4.10. These mean 
monthly gradients were computed using an accurate temperature model devel- 
oped by Dempsey (Ref. 55, 56) which has been verified with actual data from 
several sources. The mean values were determined by calculating the gradient 
for every three hours throughout the day and night for over a year's time, 
and then conputing the mean positive and negative gradients. These values 
are means , and therefore less than the maximum values (of say, 3.0 °F/in.) 
commonly used. These values may be used for design if measured data are 
not available. Results show that the thermal gradients can be linearly 
interpolated for slabs between 8 and 12 ins. (203-304 mm ) in thickness 
and also for a greater thickness. 

4.5.5 PCC Fatigue Computation . A fatigue analysis procedure was developed 
based upon the results of previous sections to provide a method of estimation 
of traffic "damage" that could result in cracking of the slabs. The basic 
fatigue design philosophy for zero-maintenance plain jointed pavements is that 



148 



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linear cracking must be prevented. This is possible through direct considera- 
tion of traffic loadings, slab curling, joint spacing, and foundation support. 
The PCC slab is subjected to many applications of heavy traffic loads. At 
the same time, it is also experiencing stresses due to temperature gradients 
which have been shown to have significant effect. Curling of the slab also 
results in "gaps" between the slab and the subbase which increase the stress 
under load. 

The major steps in the fatigue analysis are as follows: 

(1) Determine axle applications in each single and tandem axle load 
group. 

(2) Select trial slab/subbase structure, transverse joint spacing, PCC 
strength, thermal gradients, and other required factors. 

(3) Compute fatigue damage occurring at the slab edge for a given 
month, both day and night using the Miner's accumulative damage 
hypothesis (Ref. 30) and sum monthly over the entire design 
period. 

k=p j=2 i=m n. ., 
DAMAGE = I Z Z ^£- (4.13) 

k=l j=l i=l ijk 

where DAMAGE = total accumulated fatigue damage over the design period 

occurring at the slab edge 

n... = number of applied axle load applications of i magnitude 

1 J K 

j_ L. 

over day or night for the k month 
N. •■ = number of allowable axle load applications of i magni- 

1 J K 

tude over day or night for the k month determined from 

PCC fatigue curve 

i = a counter for magnitude of axle load, both single and tan- 
dem axle 

j = a counter for day and night (j=l day and j=2 night) 

k = a counter for months over the design period 

m = total number of single and tandem axle load groups 

p = total number of months in the design period 



151 



The fatigue damage is computed at the slab longitudinal edge because results 
from field observations of many jointed concrete pavements (Sec. 2.2 and 
4.3) and analytical fatigue analysis (Sec. 4.3) definitely showed this to 
be the critical point where cracking initiates. 
Applied traffic, n. .. . The n... is computed using the traffic data for the 

I J K I J K 

month under consideration. It is computed using the following expression: 

n... = (ADTm)(T/100)(DD/100)(LD/100)(A)(30)(P/100) (4.14) 
1JK (C/100)(DN/100)(TF/100)(C0N/100) 

where 

ADTm = average daily traffic at the end of the specific month 
under consideration 

T = percent trucks of ADT 

DD = percent traffic in direction of design lane 

LD = land distribution factor, percent trucks in design lane in 
one direction 

A = mean number of axles per truck 

P = percent axles in i load group 

C = percent of total axles in the lane that are within 6 in. 
(152 mm) of edge 

DN = percent of trucks during day or night 

TF = factor to either increase of decrease truck volume for the 
specific month 

CON = 1 for single axles, 2 for tandem axles 

Allowable Traffic, N. ... The N... is computed from PCC fatigue considerations. 

1 J K 1 J K 

First the total stress occurring at the edge of the slab for a given axle load 
is computed considering both traffic load and slab curling for the given 
month for either day or night conditions as illustrated in Figure 4.40. The 
stress is computed for edge loading of both single and tandem axles using 
models developed from the finite element program that realistically considers 



152 



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153 



both load stress and slab curling (as discussed in Section 4.1). The stress 

models were derived using multiple stepwise regression techniques from a 

factorial of data obtained from the finite element program over a wide 
range of the design variables. 

Slab Thickness (H): 8, 10, 14 ins. 

Thermal Gradient (G): -1.5, 0.0, +3.0°F/in. 

Foundation Modulus (k): 50, 200, 500 pci 

Slab Length (L): 15, 20, 25, 30 ft. 

Erodability (ES): 0, 12, 36, 60 ins. 

along slab edge 

The finite element computed stresses for all possible combinations of these 
factors are given in Tables 43-45. Individual equations were derived for 
load stress (STRL) and for thermal gradient curl stress (STRC) without 
traffic load. Results showed that the load and curl stresses could not be 
directly added to obtain the combined overall resulting stress (STRT). 
Therefore an adjustment factor, R, was derived to adjust the curl stress 
so that is could be added directly to obtain the correct total combined 
stress (STRT) which is used in the fatigue analysis. The total stress at 
the bottom of the slab edge with the load located at the edge is computed 
as follows: 



where 



STRT = STRL + (R)STRC (4.15) 



STRT = total resultant stress in the longitudinal direction at the 

bottom of the PCC slab edqe when the wheel load is located 

at the slab edge (load is single axle or tandem axle), see 
Figure 4.40. 

STRL = stress at bottom of PCC slab edge when load is located at 
slab edge (no thermal curling stress) 



154 



STRC = stress at bottom of PCC slab edge caused by curling of 
slab due to thermal gradient (no traffic load) 

R = adjustment factor for STRC so that it can be combined with 
STRL to give correct STRT 

The R ranges from about 0.8 to 1.5 depending on slab/foundation conditions. 
The regression equations determined for these stresses are as follows: 

Load Stress for single axle load : 

STRL = [L0AD/(18.0H 2 )][17. 35783 + 0.7801 ES — 0.05388H 3 /k + 7.41722 log 1Q (H 3 /k)] 

(4.16) 
Load stress for tandem axle load : 

STRL = [L0AD/(36.0H 2 )][14. 09599 + 0.10522ES - 0.09886H 3 /k + 6.2339 log 1Q (H 3 /k) 

+ 1.95266 (log 10 (H 3 /k)) 2 - 0.71963 log ]0 (ES + 1.0)] (4.17) 
Curl stress : 
STRC = [(G)(ET)/(5 x 10" 6 )][0. 00671 2k + 79.07391 log 1Q k + 11.72690L 

- 0.00720kL - 3.22139L log 1Q k - 0.06883LES - 0.59539ES log 1Q k 

- 204.39477H/k - 38.08854L/H - 8.36842H log 1Q k + 0.071 51 ESH 

+ 0.95691LES log 1Q k + 0.20845LH log 1Q k + 0.00058LHk - 0.00201LES log 1Q k] 

(4.18) 
R adjustment factor : 

R = 0.48039 + 0.01401H - 0.00427ES - 0.272786 - 0.00403L + 0.19508 log, k 

i uy 

+ 0.45187G log 1Q H - 0.00532G 2 + 0.01246GL - 0.00622GL log-^k 
+ 8.7872 log ]0 (H 3 /k)/H 2 + 0.00104GES - 0.11846G log 1Q (H 3 /k) 



+ 0.07001 log 1Q (ES + 1.0) - 0.01331G (4.19) 



where 



LOAD = total load on single or tandem axle, pounds 
H = PCC slab thickness, inches 
G = thermal gradient through slab, °F/in. 



155 



k = modulus of foundation support (top of subbase, pci) 
L = slab length, ft. 
ES = erodability of support along slab edge, inches 
ET = thermal coefficient of contraction of PCC/°F 
Standard estimate of error based on 426 data points: 
Single axle of STRL = 7.65 psi 
Tandem axle of STRL = 4.59 psi 
Curling stress STRC = 15.36 psi 
Total combined stress STRT = 18.97 psi 

Further verification of these equations was made by comparing the edge 
stress computed by the finite element program with the combined stress (STRT) 
computed using Equations 4.15 - 4.19 for 48 randomly selected slabs with 
varying H, G, k, L, ES and axle loads. Results of this comparison are 
shown in Figure 4.41. The standard error is 29.75 psi which, as expected, 
is higher than the 18.97 psi obtained using the data for which it was derived. 
However, the standard error is believed acceptable when compared to the 
other uncertainties involved. 

A computer program, called JCP-1 , was developed to compute the accumu- 
lated fatigue damage according to Equation 4.15 over the design life of the 
pavement. This data can be used to evaluate and design a plain jointed 
concrete pavement considering fatigue damage. 

4.6 LIMITING FATIGUE CONSUMPTION 

The fatigue analysis that has been developed is quite comprehensive 
in that it considers directly the effects of traffic loadings, thermal gra- 
dient through slab, joint spacing, loss of foundations support (i.e. pumping), 
and the increase in PCC strength over time. However, there are several 



156 



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factors that are not considered due to insufficient information. One of 
the most important factors may be the use of PCC fatigue curves obtained 
from small beams to estimate the fatigue life prediction of large fully 
supported pavement slabs. Traffic loading conditions also differ con- 
siderably between laboratory and field. Other inadequacies could be cited, 
however, the point to be made is that the final accumulated fatigue "damage" 
based on Eq. 4.15 computed for a pavement slab must be correlated with 
measured slab cracking before a limiting fatigue consumption can be selected 
with confidence for design. 

According to the Miner hypothesis, a material should fracture when the 
accumulated "damage" equals 1.0. Even if the Miner hypothesis were exactly 
correct, variability of material strengths, loads, and other properties would 
cause a variation in accumulated "damage" ranging from much less than 1.0 to 

much greater. For example, the PCC concrete laboratory data shown in 

-4 
Figure 4.35 shows a computed "damage" ranging from roughly 5 x 10 to 

more than 1000 based on the mean fatigue curve. 

Assuming that the Miner hypothesis is exactly correct, the use of a 
computed "damage" of 1.0 (or 100 percent) in design would result in a 50 
percent chance of fatigue failure if no safety factors were applied to any 
of the input parameters. Hence a design "damage" of much less than 1.0 is 
essential to provide for a very low probability of fatigue cracking (or high 
design reliability). This limiting value may be different depending on the 
type of pavement under design (i.e. low volume or expressway), and in the pre- 
sent application it is for high volume heavily trafficked expressways. 

The appropriate accumulated fatigue limiting "damage" for use in zero- 
maintenance design was determined as follows: 

(1) Field data were collected from three sources: 



158 



(a) AASHO Road Test Sections on Loops 4, 5, and 6 which were under 
traffic for 2 years (1958-1960). Sections having excessive 
pumping were excluded (i.e. pumping index greater than 40 
cubic ins. /in. of slab). A total of 28 sections were in- 
cluded ranging in slab thickness of 8-12.5 ins. 

(b) AASHO Road Test Sections that were left in-service after the 
end of the test and became incorporated as part of 1-80. These 
sections were under regular mixed traffic from 1962 to 1974. 
Extensive measurements of distress were made periodically by 
the Illinois Department of Transportation. A total of 25 
sections were included ranging in slab thickness from 8 to 
12.5 ins. 

(c) Twelve in-service projects located in the U.S. and Canada: New 
Jersey, Arizona, California (2), Michigan, Utah, Georgia, Colorado 
(2), Washington, Texas, and Ontario, Canada. These projects 

were high traffic volume freeways ranging in age from 6 to 34 
years. 

(2) The total accumulated fatigue "damage" from the opening of each 
project to traffic until the date of survey was computed using the computer 
program JCP-1 (or Eq. 4.15). The heaviest traveled truck lane was always 
used in the computations. A summary of this data along with other information 
for each project is given in Table 4.11. 

(3) A plot of cracking index (ft./lOOO ft. ) versus total accumulated 

fatigue "damage" is shown in Figure 4.42. Results show that as computed 

-4 
fatigue "damage" increases to more than approximately 1x10 the amount 

of observed transverse cracking occurring increases. Those projects having 

accumulated fatigue "damage" greater than 10" generally have large amounts 

of cracking. The abrupt increase of cracking after a long period of no 

cracking is typical of concrete pavement performance. 

The relationship between slab cracking and computed fatigue "damage" 
indicates the following: 

(a) A correlation exists between transverse slab cracking and computed 
fatigue "damage" in the slab. Fatigue damage is computed at the longitudinal 



159 





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edge and hence should relate directly to transverse cracking. This high 
correlation provides reasonable verification of certain aspects of the 
fatigue computation procedure. 

(b) A limiting fatigue "damage" can be selected for design depending 
on the amount of cracking that is acceptable in the pavement. 

The fatigue "damage" design limit for zero-maintenance design must be 

chosen such that transverse cracking is wery unlikely to occur. All 

-4 
projects having less than 10 damage either did not exhibit any cracking 

or only a yery minor amount. The amounts of probable transverse cracking 

for various accumulated "damage" determined from the curve fit through the 

data is as follows: 



(Example for L=15 ft. ) 
Percent Slabs Cracked 





2 

17 

30 

45 

-4 
A limiting "damage" value for zero-maintenance design of 10 is chosen to 

provide a high reliability that the pavement slabs will not exhibit cracking. 



Fatigue 


Damage 


(ft./lOOO ft. 2 ) 
Cracking Index 


10" 


■5 





10" 


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10" 


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1.5 


1 




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10 




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30 



164 



CHAPTER 5 
DESIGN OF JOINTS, SHOULDERS AND SUBSURFACE DRAINAGE 

All components of the pavement must be designed as part of a system. 
The structural design of the slab and subbase should be coordinated with 
the joint, shoulder, and subsurface drainage design. These components 
must compliment the structural design so that the pavement can withstand 
both heavy traffic and severe environmental factors without the occurrence 
of distress. Many conventional pavement have adequate structural design, 
but due to inadequate consideration of joints, shoulders, and/or subsurface 
drainage premature distress and reduced maintenance-free life has occurred. 
Development of design criteria are provided for joints, shoulders, and sub- 
surface drainage. 

5.1 JOINTS 

Joints are placed in plain jointed concrete pavements to control cracking 
and also for construction purposes. The major distresses related to joints as 
identified in Chapter 2 are (1) faulting (2) sealant damage and infiltration 
of incompressibles and moisture, and (3) increase in various distresses when 
joint spacing is increased. These distresses must be prevented to provide a 
maintenance-free pavement. Recommendations for minimizing their occurrence 
are presented. Excellent joint design and construction recommendations for 
other factors are provided in References 69 and 77. 

5.1.1 Joint Faulting . The faulting of transverse joints in plain 
jointed concrete pavements is a \/ery serious distress which leads to main- 
tenance when roughness becomes excessive. Most faulting problems have 
occurred on non-doweled pavements but some has also occurred on doweled 
pavements. Two major field studies have been conducted to determine the 



165 



cause of faulting in non-doweled transverse contraction joints. Conclusions 

from the California study are as follows: 

"1. Faulting of PCC pavement joints is caused by an accumulation 
of build up of loose material under the slabs near the joints. This 
accumulation may occur only under the approach slab, or it may be a 
differential build up under both slabs with the thicker layer under 
the approach slab. 

2. The build up is caused by violent water action on available 
loose or erodible materials that are beneath or adjacent to the slabs 
The water is moved backward (and probably transversely) by the fast 
depression of the curled or warped slabs under heavy wheel loads and by 
the suction caused by the release of load on the approach slab, which 
erodes and transports any loose materials. 

3. The major sources of the buildup are the untreated shoulder ma- 
terial and the surface layer of the cement-treated base. Minor amounts 
may come from abrasion of the concrete joint interface and from ma- 
terial on the pavement surface moving downward through the joints." 
(Ref. 75). 

The Georgia study concluded similarly: 

"The faulting occurring on the Georgia Interstate System is caused 
by the erosional effects produced by a combination of free water under 
the slab and the repeated passage of heavy axle loads across the joints. 
There is a loss of fine material from the top of the base on both the 
leave and approach side of the joint but this loss is more serious at 
the leave side. Some material is being redeposited under the approach 
side." (Ref. 74). 

The degree of load transfer across a transverse joint and the ability 
of the joint to maintain sufficient load transfer is related to joint faulting 
If a joint could maintain a high load transfer, joint deflection would be mini- 
mized and there would be no differential joint deflection between slabs, which 
would minimize backward pumping of fines as a load passed over the joint. 
Joint load transfer efficiency or effectiveness has been defined two ways. 
One definition is Eq. 4.2 and the other definition that is more generally used 

is as follows: 

2 d 
Joint Effectiveness (%) = ^ — ^ — 100 (5.1) 

where d, = deflection of unloaded slab 
d 2 = deflection of loaded slab 

166 



Joint effectiveness or efficieicy depends on several factors including: 
joint opening, number of load applications, foundation support, wheel load 
applications, foundation support, wheel load, aggregate particle annularity, 
and the presence of mechanical load transfer devices (Refs. 71, 72, 77). 
Results obtained from laboratory, field and analytical studies are pre- 
sented to show the effects of these factors. 

1. Joint opening: The effectiveness of load transfer is significantly 
lost as joint opening increases as shown in Figure 5.1. The rate of loss 
is much greater when the joint contains no mechanical load transfer (i.e. 
dowels) as will be subsequently shown. Joint opening depends on several 
factors such as joint spacing, frictional resistance of subbase, PCC pro- 
perties including thermal contraction and drying shrinkage, and moisture 
content and temperature of the slab. Variation in joint opening from joint 
to joint in a given pavement is high with a coefficient of variation of 40 
percent computed on one project (Ref. 73). Despite the many complexities 
involved, mean joint opening over a yearly or daily time interval can be 
computed approximately using the following expression: 

AL = CL[aAT + e] (5.2) 



where 



AL = joint opening caused by temperature change AT 
and drying shrinkage of PCC 

a = thermal coefficient of contraction of PCC (/°F) 
(generally 5-6 x 10" 6 /°F) (9-10.8 x 10" 6 /°C) 

e = drying shrinkage coefficient of PCC (approximately 0.50 
to 2.50 x 10" 4 in/in.) (cm/cm) 

L = joint spacing (ins.) 

AT = temperature range (for design use temperature at placement 
minus lowest mean minimum monthly temperature) 



167 



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(9 in. slab, 6 in. gravel subbase, k = 145 psi) 



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Loading Cycles, 100 000 



8 9 10 



(b) Influence of joint opening on effectiveness 
(9 in. slab., cement stabilized subbase, 
k = 542 pci) 



Figure 5.1. Joint load effectiveness for gravel and cement treated 
subbase for various joint openings (Ref. 71). 



168 



C = adjustment factor due to subbase/slab frictional 

restraint (0.65 for stabilized subbase, 0.80 for granular 
subbase. ) 

The recommended adjustment factor, C, and the drying shrinkage coef- 
ficient suggested were computed from limited field data from Utah (Ref. 73), 
Florida (Ref. 72), Michigan (Ref. 70), and California (Ref. 101). 

Using typical values for slab and temperature drop, maximum joint 
openings are computed in Table 5.1. Values of joint openings during ten 
years of performance in seven states with joint spacing of 10 to 25 ft. (3-7. 6m) 
range within the values shown in Table 5.1 (Ref. 70). However progressive 
opening up of joints due to infiltration of incompressibles would increase 
the joint width beyond these values. 

Joint faulting is related to joint opening and joint effectiveness. Data 
from the Minnesota road test, for example, show that joint faulting of plain 
concrete pavements without dowels increases as joint spacing (and opening) in- 
creases from 15 to 30 ft. (4.6-9.1m) (Ref. 70). Joint opening in a given cli- 
mate can be limited most easily through control of joint spacing and use of a 
stabilized subbase. Thus, if a maximum joint opening of 0.04 ins. (1.0mm) were 
considered as the limiting criteria, the yearly range in average daily tempera- 
ture is 60°F (33°C) for a given region, and a stabilized subbase is used, the 
maximum allowable joint spacing computed from Eq. 5.2 is 12 ft. (3.7m). Of 
course not all joints open the same amount and this contributes to considerable 
variability in joint faulting. 

2. Load applications: Joint effectiveness decreases as load applications 
are applied to a given joint in both laboratory and field studies (Refs. 71, 72) 
Even under heavy loadings most pavements do not fault until 5-10 years of ser- 
vice. The rate of decrease depends on many factors including joint opening, 
slab thickness, foundation support, load magnitude, mechanical load transfer, 



169 



Table 5.1. Computed maximum joint openings using Eq. 5.2 for a 
temperature drop of 60°F and drying shrinkage. 









Joint 


Opening 


- ins. 








Stabi 


lized 


Subbase 




Granul 


ar 


Subbase 


Joint 
Spacinq-ft. 


Temp. 




Temp. & 
Shrinkage 


i 


Temp. 




Temp . & 
Shrinkage 


10 


.026 




.034 




.032 




.041 


15 


.039 




.050 




.047 




.062 


20 


.051 




.067 




.063 




.082 


25 


.064 




.084 




.079 




.103 


30 


.077 




.101 




.095 




.124 



5.5 x 10" 6 /°F 



AT = 60°F 



e = 1.0 X 10 



-4 



170 



and type of aggregate (See Figures 5.1 and 5.2). 

3. Slab Thickness and Foundation Support: Increased slab thickness and 
foundation support decreases the loss of joint effectiveness with repeated 
load applications as shown in Figures 5.1 and 5.2. This may be due to the 
reduced unit vertical shear across the joir.'. face as the slab thickness and 
k-value are increased. The corner deflection under a load changes with slab 
thickness and foundation support as shown in Figure 5.3. As the k-value de- 
creases below 200 pci (54 MN/m ) the deflection increases rapidly. The less 
the deflection the less unit shear that must be transferred across the joint 
for a given load, and thus over a number of repeated load applications the 
joint load effectiveness remains higher. 

4. Load Magnitude: As load magnitude increases the joint effectiveness 
decreases as shown in Figure 5.2 because deflection and unit shear transferred 
across the joint for every load application is increased. 

5. Mechanical Load Transfer: Dowels improve the joint effectiveness 
significantly for relatively wide open joints as shown by some 12 year 
field results from Florida in Figure 5.4. Data from a recent Georgia study 
also indicate higher load effectiveness for doweled joints, particularly 
during cold weather (Ref. 78). The most significant effect of dowels is in 
reducing the faulting at transverse joints. Considerable evidence can be 
cited to show that the use of dowels causes a significant reduction in faulting 
over a similar joint without dowels. The most recent and comprehensive study 
was conducted by the Florida DOT where faulting was measured at seven test sites 
over a 12 year period. Both doweled and non-doweled plain jointed concrete 
pavements were located at most of the sites so that a direct comparison could 

be made. Histograms showing measured faulting for some of the test sites are 
given in Figures 5.5-5.6. Conclusions reached by Stelzenmuller, Smith, and 
Larsen are: 

171 



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(b) Influence of wheel load on effectiveness 



Figure 5.2. Joint load effectiveness (Ref. 71) 



172 



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Test Sites 1-7 
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Pavement Age -4 to 6 Years 




fflfflpfl P P p p p p 



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Pavement Age - 4 Years 







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0.02 0.04 0.06 0.08 0.10 

Joint Faulting (inches) 



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0.14 



Comparison of Joint Faulting in Florida for Doweled 
arid Non-doweled Joints. 



175 



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176 



"It was found that doweled joints show less faulting than non- 
doweled joints and that dowels increase the effective load transfer 
when properly designed and installed. Aggregate interlock provides 
40 percent or more load transfer during warm weather, but frequently 
less than 40 percent during cold weather. Dowels retard joint fault- 
ing better than aggregate interlock." (Ref. 72) 

Results from the Michigan, Minnesota, and Missouri road tests (Ref. 70) indi- 
cate that doweled joints definitely fault less than non-doweled joints. Van 
Breemen concludes that for heavily trafficked pavements in New Jersey, load 
transfer devices are needed at transverse joints to avoid the faulting of 
those that open excessive amounts (Ref. 51). Studies in Georgia also show 
that dowels are effective in reducing faulting (Ref. 74). 

Even though the use of dowels have definitely reduced faulting, there 
are several projects that had doweled joints, yet still showed serious fault- 
ing. Hence dowels are definitely not a panacea for prevention of faulting. 
A good example is the AASHO Road Test sections of plain jointed concrete 
pavement left in service on 1-80 for 12 years as shown in Figure 5.7. The 
8 in. (203mm) slabs showed very serious faulting, but faulting decreased with 
increased slab thickness and dowel diameter. The bearing stress on the con- 
crete was computed using the Friberg analysis method (Ref. 76) and a plot of 
bearing stress caused by an 18-kip (80 kN) single axle load at the joint 
versus joint opening for each of the four slabs is shown in Figure 5.8. 
Mean joint faulting versus dowel concrete bearing stress is shown in Figure 
5.9. There is a high correlation between bearing stress and faulting. These 
joints have been subjected to approximately 13-19 million 18-kip (80 kN) ESAL 
and there has been some pumping of the subbase, 

5. Aggregates: Joint effectiveness is maintained at a higher level when 
the aggregate hardness and angularity is increased (Refs. 71 and 72). The 
effects are most significant after many load applications. 



177 




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Figure 5.8. Computed Concrete Bearing Stress Under Dowels for JCP-1 to 25, 



179 



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180 



Griffin concluded that prevention of faulting can be accomplished by 

(1) excluding free water from the subbase in the joint area, (2) selecting 
porous subbase materials or those not affected by erosion, and (3) making 
the load transfer device adequate (Ref. 79). Recommendations of Spellman, 
Woodstrom, and Neal from the California study include: (1) use of dowels, 

(2) shorter joint spacing (6-8 ft. ) (1.8-2. 4m), (3) increased thickness 
of slab, (4) minimize water infiltration, (5) subsurface drainage, and 
(6) use a erosion-resistant subbase and shoulder materials (Ref. 75). 

The following conclusions are made with regard to prevention of trans- 
verse joint faulting for pavements subjected to heavy track traffic: 

1. A high degree of joint load transfer effectivness must be maintained 
throughout the design life through either aggregate interlock or both aggre- 
gate interlock and mechanical load transfer. The joint load transfer effect- 
iveness must be maintained throughout the year, particularly in cold weather. 
Thus, maintaining adequate joint effectivness in colder climates using only 
aggregate interlock is not believed possible under heavy traffic. 

2. Dowel bars of adequate size and spacing will greatly help to maintain 
joint effectiveness and reduce faulting even under heavy traffic if slab thick- 
ness and subbase support are adequate to limit concrete bearing stress. 

3. Relatively short joint spacing must be used so that seasonal and 
daily joint opening can be kept within acceptable limits to maintain aggre- 
gate interlock. Based on laboratory and field studies an acceptable limit is 
approximately 0.03 in. (0.8mm). However, due to the large variation in joint 
openings there will be a large proportion of joints with openings greater than 
this limit even if the design opening is less than 0.03 in. Thus, aggregate 
interlock will be lost and faulting may occur at these joints if dowels are 
not provided. 

181 



4. An open graded drainage layer beneath the slab in wet climates will 
minimize free water and high water pressure beneath the slab. If the subbase 
is dense graded, it must be erosion resistant. 

5. Joints should be adequately sealed to reduce water and incompressibles 
infiltration into the joint. This will reduce free water under the slab and 
prevent progressive joint opening. 

6. Increased PCC slab thickness will increase the area of effective 
aggregate interlock and reduce joint deflection under load thereby improving 
long term joint effectiveness. It also will reduce concrete bearing stresses 
caused by dowels. 

7. Increased slab foundation support will reduce joint deflection and 
hence increase long term joint effectivensss and reduce concrete bearing stress 
(minimum recommended is 200 pci (54 MN/m ). 

8. Provision of full depth PCC or asphalt concrete shoulders will mini- 
mize availability of pumpable material. PCC shoulders tied to the traffic 
lane will provide increased load transfer along the longitudinal joint 

and hence reduced deflection. 

5.1.2 Joint Sealant Damage . The major problems commonly associated with 
the deterioration or damage of joint sealants include: (1) increased water 
infiltration causing significant free water beneath the slab, and subsequent 
pumping and other moisture problems, and (2) infiltration of incompressibles 
which may result in progressive joint opening (pavement growth), spalling, 
blowups and loss of load transfer. The interviews and field surveys of plain 
jointed concrete pavements, however, indicated that blowups and joint spalling 
rarely occured. Only two projects were observed to exhibit pavement growth 
(in New Jersey and Arizona) due to infiltration of incompressibles. Most states 
visited do not reseal joints on heavily trafficked pavement as evidenced by the 
fact that only a few of the pavements surveyed had resealed joints over time 



182 



periods of 6 to 30 years. Thus, it appears that some engineers are not much 
concerned about the deterioration of joint seals for short jointed plain con- 
crete pavements. There are certainly many, however, that do feel that the 
joint seals are important. The degree to which joint faulting occurred be- 
cause of infiltration of moisture and incompressibles is believed to be sig- 
cant on several of the projects visited. Recent studies have shown that with 
current sealant materials and practices it is impossible to keep moisture out 
of the pavement section (Refs. 82-86). If an open graded drainage layer is 
provided in "wet" areas and a non-erodable subbase is provided in "dry" areas 
the effect of moisture on faulting will be minimized. Thus, it is definitely 
desirable to provide a sealant that will at least minimize the infiltration of 
moisture and also prevent the intrusion of incompressibles. 

Only the most durable sealants should be used for zero-maintenance design 
and the expected maximum joint opening and allowable sealant extensibility should 
be considered when selecting the sealant. Two general types of sealants are 
available: liquid and preformed. Field studies have shown that hot applied 
bituminous liquid sealants have a field life of only a few years and thus would 
not be suitable for long term performance (Refs. 2, 84, 87, 88). For example, 
a recent study by McBride and Decker (Ref. 88) shows the bituminous sealant 
had failed and significant infiltration of incompressibles occurred on pave- 
ments having a random joint spacing of 12-18 ft. (3. 6-5. 5m) and varying in 
age from 1-1/2 to 10 years (See Figure 2.15). However, the newer elastomer 
types (polysulfides, urethanes, polyvinyl chloride, etc.) may be capable of 
providing longer service lives (some companies are now providing a warranty 
for 10 years on these seals). Preformed sealants are available that have been 
shown to give 10+ years of acceptable service even on long jointed concrete 
pavements under heavy traffic. Figure 5.1Ca is a photo of a preformed seal 
in service on Highway 401 Toronto for 12 years (joint spacing is 56 ft. (17m), 



183 



(a) Hwy 401, Toronto, Ontario, 12 years old, 
joint spacing 56 ft, ADT 187,000 




(b) Hwy 27, Toronto, Ontario, 6 years old, 
Joint spacing 12-19 ft, ADT 80,000 



Figure 5.10. Neopreme Preformed Compression Sealant In-Service 



184 



and Figure 5.10b shows a preformed seal on JCP-37 in Toronto in service for 
6 years (joint spacing 12-19 ft. (3. 6-5. 8m), with very few signs of distress. 
Preformed compression sealant was the type most recommended to achieve the 
required long term performance needed. Preformed seals have been observed 
to do an excellent job of keeping out incompressibles over long time inter- 
vals, but are not completely water tight seals. 

The liquid or preformed seal must be designed to accommodate joint move- 
ment. The elastomers have an expansion-compression range of about + 20 per- 
cent at temperatures from -40°F to +180°F (-40 to 82°C), The preformed 
sealants are designed so that the seal will always be in compression. The 
preformed compression sealants should be selected so that it will always be 
compressed at least 20 percent in the sawed joint. The maximum allowed com- 
pression of the seal is 50 percent before a rubber on rubber situation is 
reached. Thus the seal working range is 20 to 50 percent (Refs. 88, 89, 90). 

The required joint sawed width for a liquid sealant can be determined 

as follows: 

Joint spacing = 15 ft. (4.6m) 

Design temperature range = 100°F (55°C) (temperature at placement 
minus lowest mean minimum monthly temperature) 

Stabilized subbase 

Maximum allowed extension of sealant = 20% 

Design joint movement (Eq. 5.2): 

AL = 0.65 x 10 ft. x 12 j± [5.5 x 10' 6 /°F x 100°F + 1.0 x 10" 4 ] 
= 0.076 ins. (1.9mm) 

Min. joint sawed width = * 2n = 0.38 in. (9.7mm) 
Use reservior width = 0.5 in. (12.7mm) 



185 



The size of preformed seal and joint sawed width can be determined as 

follows: 

AL = 0.076 in. (1 .9mm) 

Try 7/16 in. (11mm) width seal in a 1/4 in. (6.4mm) sawed 
joint and install in summer (joint would not be further 
compressed) . 

4375 - 25 
Check maximum Compression: — — Q 4375' — x 100 = 43% < 50% ok 

Check minimum Compression: — n~4i7R '' x ^ 00 

= 25% > 20% ok 
Working range of seal = 25 - 43% of uncompressed seal 

In conclusion, while there is not yet a sealant available that can be 
guaranteed to last for 20 years, the minimum desired zero-maintenance design 
period, there are sealants that will definitely perform satisfactory for more 
than half of this time period. Considering the numerous heavily trafficked 
plain jointed concrete pavements that never receive joint maintenance, and 
the structural design being proposed for zero-maintenance pavements (thick 
slabs, subsurface drainage, dowel bars, stabilized subbase, stabilized 
shoulders, and short joint spacing), it is fairly certain that joint sealant 
maintenance will not be necessary for the rest of the design period should 
some of the joint sealant fail after 10-15 years. 

5.13 Transverse Joint Spacing . This factor has a very significant 
effect on pavement performance. Decreasing transverse joint spacing has the 
following beneficial effects: 

1. decreases thermal curl stress (Fig. 4.31) 

2. decreases transverse cracking (Fig. 4.30) 

3. decreases curlina of slab upward at joint 

4. decreases joint spall ing (Fig. 5.11) 

5. decreases seasonal and daily joint opening, and thus: 

- increases joint load transfer effectiveness. 



186 





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1G7 



- reduces sealant extension 

6. For spacings less than about 20 ft. (6.1m) increased slab 
thickness is beneficial in decreasing total stress from load 
and temperature gradients (Fig. 5.12). 

Of course, to minimize construction cost it is desirable to increase the 
joint spacing, however, maintenance-free performance is of prime concern here. 
Based upon these considerations, the recommended maximum joint spacing for 
zero-maintenance pavements is 20 ft. but it is highly recommended to limit spacing 
to about 15-17 ft if dowels are used, and 12-15 ft. or less if they are not used. 

5.14 Joint Load Transfer Device . Many types and varieties of mechanical 
load transfer devices (LTD) and joint configurations have been tested over the 
years. Nearly all of these have given poor performance and caused serious 
joint failure. Several pavement failures caused by proprietary lugs were 
observed during the field survey. Based on extensive field experience, round 
steel dowels have become the standard and nearly the only mechanical LTD used 
in current practice. There are also significant problems associated with dowels 
as determined from the interviews, field and laboratory studies: (1) corrosion, 
(2) providing proper size and spacing for the specific application, and (3) align- 
ment. 

1. Corrosion: Although the serious consequences of dowel corrosion have 
been known for many years, almost nothing has been done to prevent it until 
just the last few years. The most significant and perhaps first work to determine 
the effects of dowel corrosion and the subsequent development of corrosion proof 
dowels was by Van Breemen of the New Jersey Department of Transportation. Van 
Breemen first reported the effects of corrosion in New Jersey in 1945 (Ref. 81) 
and a comprehensive report on field and laboratory studies was published in 
1955 (Ref. 80). Major findings from these studies are as follows: 

1 . Serious corrosion of carbon steel dowel bars occurred after only a 

few years in service. Most corrosion occurred in the joint space, but 

188 



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8 9 10 II 12 13 14 
PCC Slab Thickness , inches 



Figure 5.12. Effect of Joint Spacing and Slab Thickness of Curling 
and Combined Curling and Load on Edge Stress (computed 
with Finite Element Program). 



189 



extended for 3 inches or more on each side of the joint. 

2. The corrosion resulted in a lock-up of the joint as verified by 
joint width change measurements and dowel pull out tests. The joint lock- 
up resulted in the forced opening of transverse cracks and rupturing the slab 
reinforcement. These cracks soon spalled and faulted. Some joints also showed 
distress (spalling and faulting). 

3. Several experimental dowels to which had been applied various coat- 
ings and protection treatments were installed in pavements and evaluated after 
8 years in service. 

"1. All of the dowels partly encased in Monel tubing 
("monel" dowels) were still in practically perfect condition. 

2. All of the ordinary hot-rolled carbon steel dowels 
were rusted in various amounts, ranging from minor to serious, 
regarldess of the kind of coating. 

3. Hot-dipped galvanizing delayed but did not prevent 
rusting. 

4. All of the various dowel coatings (which included 
red lead, white lead, tar paint, graphite paint, transmission 
oil, cylinder stock grease, and asphaltic oils, grades MC-2, SCO, 
and CC) had deteriorated to the extent that they are now practi- 
cally worthless. Moreover, in numerous instances the coatings 
had disappeared entirely. Apparently the deterioration and 

loss has been due mainly to the action of water. 

5. The enclosure of the sliding portion of the dowel in 
a sheet metal sleeve had no apparent effect on prolonging the 
life of the coating, at least to any appreciable extent. 

6. In practically all instances, the exceptionally thick 
coatings of asphalt applied to the hot-rolled dowels on Route 
25, Newark, had disappeared entirely, despite the fact that 
sheet metal sleeves were used in conjunction with these coatings. 
In addition to the loss of the coatings, the sleeves had almost 
completely rusted away, and there was considerable rusting and 
significant loss in section of the sliding portions of the dowels, 
and at the joint space. 

7. Thick coatings apparently prevent seizure, but upon their 
disappearance there is a creation of clearance. This in turn per- 
mits the entrance of corrosive agents to the entire sliding portion 
of the dowel . 

8. The back-and-forth movement of the dowel results in a 
certain amount of abrasion. 

9. In practically all instances, the compressible material 
at the ends of the dowels was in a very wet condition. 

Judging from the conditions found, it appears probable 
that there are times, during rainy weather, when the joints are 
completely inundated and, in effect, functioning under water. 
There furthermore were strong indications in several instances 
that there was a slight vacancy under the dowel which served 

190 



as a channel to carry water clear to the end of the dowel. 
This vacancy was apparently the result of a slight slumping- 
away of the concrete, or the accumulation of a thin layer of 
water or air, underneath the dowel." (Ref. 80). 

Additional tests with stainless steel coated dowels (both stainless steel 
encased tubing and solid stainless steel coatings) showed that "no rusting 
of the stainless steel tubing has occurred nor does there appear to be rust- 
ing of any consequence between the tubes" (Ref. 80). Based upon these results 
New Jersey has specified Monel or stainless steel protection (as alternatives) 
for their dowels since 1947 and have obtained excellent results. Joint lock- 
up was virtually eliminated in these pavements (Ref. 2). The dowel bar speci- 
fication used by New Jersey is given in Figure 5.13. 

To further verify the serious problems of dowel corrosion, two joints 
were removed from concrete slabs near Ottawa, Illinois by the Illinois Depart- 
ment of Transportation (IDOT) in June, 1975 and then tested at the University 
of Illinois to determine the extent of joint lock-up. The entire full lane 
width was removed as shown in Figure 5.14 from Section 382 of the original 
AASHO Road Test (now 17 years old and under heavy traffic on 1-80) and from 
another section constructed in 1962 by IDOT. Two six foot wide sections of 
joint containing six dowels were pulled apart using hydraulic jacks and the 
axial tension load and joint opening carefully recorded. A plot of joint 
opening versus axial tensile force as is shown in Figure 5.15 for Section 382. 
A total of 30,000 lbs (133kN) was required to open the joing 0.1 inches (2.5mm) 
(approximately 5000 lbs (22kN) /dowel ) . Visual observation of the dowels showed 
significant corrosion at least 3 inches (76mm) on either side of the joint 
(Ref. 91). The other joint tested gave approximately the same results (Ref. 
91). The pavement slabs on either side of the joints were 40 ft (12.2m) in 
length and contained highly spalled and faulted cracks, and the steel rein- 
forcement had ruptured. These results closely parallel the New Jersey fund- 
ings. Approximately 5 tons/lane mile (44kN/km) of deicing slabs are used 

191 








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on this highway per year which has definitely contributed to the corrosion 
problem. 

Additional studies on dowel corrosion and pull out tests were conducted 
by New York (Ref. 92). These results showed similar findings with regard to 
corrosion and joint lock-up. Therefore it is concluded that to prevent dowel 
corrosion and subsequent joint lock-up in areas where deicing salts are used 
(and possibly other areas where corrosion occurs), dowels must be corrosion 
resistant. Long term experience in New Jersey has shown that Monel metal or 
stainless steel coatings do provide non-corrosive dowels that allow adequate 
joint movement for pavements now in service nearly 30 year. Other corrosion 
resistant coatings are available, but without long term performance data as 
for the stainless of Monel clad dowel. These methods include pretreatment 
with various plastic coatings (Ref. 93) and fiberglass dowels. 

2. Size and Spacing: The size and spacing of dowels currently used is 
based upon field, laboratory, and analystical studies. Dowels generally are 
spaced 12-15 ins. (205-381mm) and have a diameter of approximately 1/8 of the 
slab thickness. These criteria have generally been successful in preventing 
faulting but there are exceptions, particularly on heavily trafficked pavements 
(i.e. the previously discussed joints shown in Figure 5.7). The analytical 
dowel analysis and design procedure such as that of Friberg (Ref. 76) and 
ACI (Ref. 94) are based on providing sufficient strength to transfer one-half 
of the assumed pavement design wheel load across the joint with a safe bearing 
pressure between the dowel and the supporting concrete. Allowable bearing stress 
as recommended by the ACI and the computed bearing stress and mean joint fault- 
ing for the AASHO Road Test sections in-service on 1-80 previously discussed 
are given in Table 5.2. These data show that even though the allowable bear- 
ing stress was not exceeded (the allowable stress is based on static failure 
criteria) considerable faulting still occurred even though the dowels were of 

195 



Table 5.2. Joint concrete bearing stress and faulting data 



Slab Dowel Concrete Allowable Mean Joint 
Thickness Diameter Bearing * Bearing ^ Faulting 
(in) (in) Stress-psi Stress-psi (in) 



8 


1 


2715 


3000 


0.25 


9.5 


1 1/4 


1620 


2751 


0.07 


11 


1 3/8 


1230 


2625 


0.04 


12.5 


1 5/8 


840 


2376 


0.03 



* Computed using Friberg Analysis (Ref. 94) 
**Computed using f^i = 3000 psi 



196 



the recommended size and spacing. This is probably the result of a concrete 
bearing fatigue failure phenomenon. This data may indicate that for heavy 
truck traffic (i.e. approximately 1 million 18-kip (80kN) ESAL/year/lane) 
the concrete allowable bearing stresses should be further reduced as indicated 
in Table 5.2. Further research in this area is clearly needed. One final 
and important point with regard to dowels is their required spacing on 
multilane pavements. The point at which joint faulting is most serious is 
the outside corner of the outer traffic lane. This is also the point at 
which critical dowel bearing stress occurs. On multilane pavement, faulting 
is much less on interior lanes thus it is believed that dowel spacing may be 
increased in these lanes. Truck travel is generally less on the interior 
lanes and there is significant load transfer across the longitudinal joint. 

5.2 SHOULDERS 

Shoulders must be designed to provide zero-maintenance performance, since 
repair of a shoulder failure usually requires closing of the adjacent traffic 
lane. Results from field studies indicate that the only shoulder types ad- 
jacent to jointed concrete pavements that have low maintenance performance 
are PCC and full depth asphalt concrete (AC) (Refs. 2, 95, 96, 97, 98, 99). 
Full depth asphalt concrete shoulders have shown generally good performance 
in the various states, but have required longitudinal joint maintenance often 
and have exhibited some separation and cracking at the longitudinal lane shoulder 
joint in freeze areas (Ref. 95). PCC shoulders have been observed to give over 10 
years of maintenance- free performance in Illinois (with no sign of distress) 
and equal performance in other states over shorter time periods. Figure 
5.16 shows a photo of a 10 year old PCC shoulder in Illinois located 



197 




10 ft. 



Anchor 
Bars 




Crushed Stone 
Cement Agg. Subbase 



The longitudinal joint in half of each test section is 

sawed 1/2" wide by 3/4" deep and sealed with a hot-poured 

rubber-asphalt joint sealant. 

Surface is transverse broomed. Dummy groove transverse 

joints at 20' intervals. 

Rumble strips 6 1 wide @ 60' intervals. 

Anchor Bars: No. 4 hooked bolts, !C In. length turned into 
2-in. snap off expanding end anchors set into 
the edge of slab at 30 in. intervals. 



Figure 5.16. PCC shoulders on 1-80 in Illinois, 10 years old 



198 



on 1-80, and Figure 5.17 shows an 11 year old PCC shoulder on 1-74. Both 
have provided zero-maintenance performance. Recommendations from highway 
agency engineers indicated perference for PCC shoulders when the traffic 
lanes were PCC. However, full depth AC shoulders may provide maintenance- 
free performance in certain climatic regions where previous performance 
data are available. 

5.2.1. PCC Shoulder Design . PCC shoulders, when tied to the lane slab, 
have the potential to reduce edge stress. Figure 5.18 was prepared using the 
finite element program to illustrate the effect of longitudinal joint effi- 
ciency. Even the small amount of load transfer obtained from tie bars may 
reduce critical edge stress as shown. The following recommendations are 
based upon results of field studies in Illinois and other states: 

(a) Slab thickness - preferably equal to the mainline slab thickness 
at the longitudinal joint and continuing at the same thickness throughout. 
If taper is more economical, begin taper 24 inches from joint to a thickness 
of eight inches, at the shoulder edge. 

(b) Tie Bars - tie the shoulder to the mainline pavement by No. 4 or 
No. 5 deformed steel tie bars spaced at 30 inches on center. 

(c) Transverse joint - space joints identical to the traffic lanes. 

(d) Longitudinal joint - place at edge of traffic lane and seal (or 
preferably increase traffic lane width 1-2 ft. to decrease number of edge loads) 

(e) Subbase - use same subbase as placed beneath PCC mainline pave- 
ment (See subsurface drainage, Section 5.3). 

Further details on PCC shoulder design can be found in Reference 100. 

5.2.2. Asphalt Concrete Shoulder Design . The following recommendations 
are based upon results of field studies in several states: 

(a) Thickness - equal to slab thickness at the longitudinal traffic lane/ 

shoulder joint for at least 24 inches and tapering to 10 inches at shoulder edge 

(edge must be of adequate thickness to prevent failure from parked trucks. 

199 





j— No. 4 Tie bars, 30 in. lono, spaced at 30 in. 

> IQ_rt \ 

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P.C.C. 6 in. 
Open Graded Subbasef- 



Porous Granular BacktILl 



Earth 



Pipe Underdrain £^ Perf. CMP I QMP 
Granular subbase 




Longitudinal keyed joint 

Rumble Strips: 4 ft wide group ;,iqs of 
corrugations 1 in. deep 
Transverse joints spaced 10-20 ft. 



Figure 5.17. PCC shoulders on 1-74 in Illinois, 11 years old 



200 




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(b) Longitudinal joint - saw cut a 1 inch square joint and fill with 
high type joint sealant. 

(c) Subbase - use same subbase as is placed under PCC mainline pave- 
ment (See subsurface drainage, Section 5.3). 

Further information on AC shoulder design can be found in Reference 95. 

5.3 SUBSURFACE DRAINAGE 

The presence of free water beneath the PCC slab can result in several 
distresses which would limit the maintenance-free life of the pavement. 
These distresses may include cracking, faulting, pumping, frost heave, and 
durability problems in PCC and subbase. Therefore, in regions where relatively 
high annual rainfalls exist, or where significant ground water exists, consid- 
eration should be given to providing subsurface drainage systems. The major 
components of a general subsurface drainage system are shown in Figure 5.19. 
The major purpose of the subsurface drainage system is to rapidly drain the 
roadbed to reduce the periods when the structure is exposed to excess moisture. 

The general guidelines for the design of subsurface drainage systems, 
developed for the Federal Highway Administration by Cedergren (Ref. 85) are 
recommended. The basic procedure considers subsurface drainage layers as 
conveyors of water and considers in-flow rates from all significant sources. 
Seepage principles are then used to determine the permeability and thickness 
of a subsurface drainage layer that will accommodate the water flow. Hence, 
the drainage system is designed to have an out-flow rate equal to the rate 
of infiltration into the pavement during a one-hour design rainfall having a 
frequency of occurrence of one year. Theinfil tration rate of PCC is considered 
to be 0.50 to 0.67 of the total rainfall. This procedure, however, gives a 
drainage layer having a very high drainage capacity. This high capacity may 
not be required and a somewhat less porous drainage layer may be adequate, 
which would have greater stability in providing support for the PCC slab. 

202 






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Pavements located in areas where subsurface moisture is not of suf- 
ficient magnitude to provide subsurface drainage should not, however, be 
constructed in a "bathtub". The subbase should be daylighted to provide 
for some lateral drainage in these areas. The subbase should be con- 
structed with high quality asphalt or cement stabilized granular materials 
to provide an erosion proof subbase surface. Additional information on 
subsurface drainage is found in References 82, 83, 84, and 85. 



204 









CHAPTER 6 
VERIFICATION OF DESIGN 

6.1 DESIGN APPROACH AND COMPUTER PROGRAM 

The structural design of zero-maintenance plain jointed concrete 
pavement involves two independent but complimentary approaches. It is 
based upon both a serviceability/performance analysis and a PCC slab 
fatigue analysis. Both approaches have distinct capabilities needed for 
structural design. The design recommended for construction must meet the 
limiting criteria of both approaches. Structural design consists of the 
selection of PCC slab thickness and strength, subbase type and thickness, 
and joint spacing. The design must be compatible with shoulder and 
subsurface drainage design. The general approach to design is shown in 
Figure 1.2. 

The basic structural design philosophy to provide zero-maintenance 
plain jointed concrete pavements is to prevent linear (transverse) 
cracking of the slab and excessive pavement roughness caused by joint 
faulting and other factors such as slab settlement. The fatigue analysis 
provides a direct precedure that is used to prevent transverse cracking. 
The serviceability/performance analysis provides a direct procedure that 
is used to prevent the occurrence of excessive roughness as indicated by 
the serviceability index which is an estimator of the user's acceptability 
of the pavement. 

Limiting criteria have been selected for zero-maintenance design for 
fatigue damage and for serviceability index as described in Chapters 3 and 

(a) Fatigue Damage : A maximum allowable fatigue damage (or DAMAGE) 
as accumulated monthly over the entire design analysis period at the slab 



205 



-4 
edge, midway between joints, is 10 as computed by Eq. 4.13. However, 

since this value is yery small and inconvenient to use in design, it 
was multiplied by a scale factor of 10 so that the limiting value is 
100. This value was set based upon fatigue analyses of 37 in-service 
pavements ranging in age from 6 to 34 years to give a high reliability 
that linear cracking from fatigue would be prevented. The computed 
fatigue damage for each of these pavements, prior to any cracks that oc- 
curred, was greater than this specified limiting value. 

(b) Terminal Serviceability : The terminal serviceability index 
selected is 3.0. This value was set based upon observations on the 37 
in-service pavements. Use of this value along with a reduction in the 
modulus of rupture provides a high reliability that the pavement will not 
require maintenance due to excessive roughness over the design analysis 
period. 

A computer program was written to provide fatigue and serviceability 
data for use in design. The program is designated: JCP-1 (Jointed Concrete 
Pavement - 1) and is written in FORTRAN computer language for the IBM-360 
digital computer. The program can be adapted for usage on other computers 
with only minor modifications. The computer processing time for a complete 
design problem is about 12 seconds. The storage requirement for the pro- 
gram is 50,000 locations. An input guide, sample input/output, flow chart, 
and program listing is given in the Appendix. The designer must specify 
trial structural designs, determine the required inputs, run the JCP-1 com- 
puter program, and analyze the output fatigue and serviceability data. 
The program is written to analyze any number of slab thicknesses and pro- 
vide outputs for each thickness, while holding all other inputs constant. 



206 



The designer can therefore examine a range of slab thicknesses for a given 
traffic, foundation support, and environment with only one run of the 
program. 

A complete detailed example of zero-maintenance design is given in 
Chapter 6 of Volume II (Ref. 1). 

6.2 STRUCTURAL DESIGN VERIFICATION 

Complete verification of the design procedure requires construction 
of the recommended designs in various climatic regions and observation of 
their performance over the design maintenance-free life. In lieu of this 
costly and time consuming procedure, a reasonable verification can be 
obtained by comparing the design and performance of various plain jointed 
concrete projects with a zero-maintenance design of each project. The 
zero-maintenance design period would be set equal to the existing life of 
the project under consideration, the design inputs would include the as- 
built construction data and the traffic applied to the project since its 
construction. It is desirable to use projects that were not used in the 
development of the design procedure. Due to the limited funding and time 
available this was not possible. Thus, the 37 projects are used to pro- 
vide a partial verification of the procedure. The following steps were 
followed for each project: 

1. As-built construction data (material strength, joint design, 
thicknesses, etc) were obtained from the agency responsible for its construc- 
tion. Traffic data over the life of the project were also obtained (See 
Chapters 2, 3, and 4). 

2. The foundation including subbase and subgrade and the joint design 
were kept the same as the existing project, and all other necessary inputs 



207 



to the zero-maintenance design procedures determined. 

3. The required slab thickness necessary to meet the fatigue damage 
and serviceability limiting design criteria were then obtained using the 
design procedure. 

4. The new design slab thickness was compared with the actual slab 
thickness and performance of the pavement. Conclusions were made based 
upon this comparison. 

A summary of results from each of the projects is given in Table 6.1, 
and a discussion of these results for selected projects is provided. 

1. AASHO Road Test Site: Existing pavements are 16 years old with 
slab thicknesses of 8, 9.5, 11, and 12.5 ins. (203, 241, 279, 318 mm) and 
a transverse doweled joint spacing of 15 ft. (4.6 m). A summary of 
significant performance data are as follows: 



Slab 


Crackir 
(ft/1000 

20 


ig o 


Joint 
Faulting 

0.25 in. 


Se 


Final 
rviceabil ity 
Index 


8 in. 




1.7 


9.5 


3 




0.07 






3.1 


11 







0.04 






3.3 


12.5 







0.03 






3.6 



The 8 and 9-1/2 inch (293, 241 mm) slabs received patching and crack 
repair. The thicker slabs did not require maintenance with the exception 
that some patching was also needed due to joint spalling caused by "D" 
cracking. The thickness of slab provided by the zero-maintenance design 
was 10.7 ins. (272 mm) with fatigue controll ing the thickness. This pavement 
constructed over the existing foundation and having similar materials and 
joint design would not crack or fault and the serviceability index is 



208 



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above 3.0 over the 16 year period. In addition, subsurface drainage is 
recommended since the project is located in a wet-freeze climate. Figure 
6.1 shows slab cracking vs. slab thickness. Full depth asphalt or con- 
crete shoulders are also recommended. Thus, the new design would be 
expected to provide maintenance-free performance over the 16 year 
period. A higher quality PCC must be provided, however, to prevent 
the spall ing caused by "D" cracking. 

2. JCP-27 and 28, California: These pavements both have 8 in. (203 
mm) PCC slabs over 4 ins. (102 mm) of cement treated subbases, non-doweled 
joint spacing of 15 ft. (4.6 m), and they are both 20 years old. Both 
projects have significant faulting and some transverse cracking. JCP-27 
has received some patching and crack filling maintenance, but JCP-28 is 
maintenance-free. However, both may require maintenance soon due to the 
joint faulting condition. It is interesting to note that JCP-28 was 
subjected to the heaviest traffic of any project surveyed (39 million 18 
kip ESAL). The zero-maintenance design slab thickness is 10.4 and 10.8 
ins. (264, 274 mm) for JCP-27 and 28, respectively. This increase in slab 
thickness would reduce joint faulting and transverse cracking on both 
projects, and should definitely provide maintenance-free performance 

over the 20 year period. 

3. JCP-32, New Jersey: This project is 25 years old and its design 
consists of 10 in. (254 mm) PCC slab, 12 ins. (305 mm) granular non-pumping 
subbase, 15 ft. (4.6 m) transverse non-doweled joint spacing, and very 
high quality PCC. It has been subjected to very heavy traffic (i.e., 

36 million 18-kip ESAL). The pavement has significant joint faulting 
but no cracking and the project has never received maintenance. The 



210 



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211 



zero-maintenance design slab thickness is 12.1 in. (307 mm) with servicea- 
bility criteria controlling the design. Dowel bars and subsurface drainage 
would be recommended for this project since it is located in a wet-freeze 
climate to prevent faulting. This thicker slab design, and other recom- 
mendations would be expected to provide maintenance-free performance 
over the 25 year design period. 

4. JCP-34 and 35, Colorado: These two 10 year old projects have 
similar designs of 8 ins. (203 mm) PCC slab, granular subbase and subgrade, 
and non-doweled transverse joints. However, JCP-34 has a longer joint 
spacing (12-19 ft. [3.7-5.8 m]) than JCP-35 (15 ft. [9.6 m]). JCP-34 has 
considerable transverse cracking and some corner cracking, .d also 
significant joint faulting and has received some maintenance. JCP-35 
does not have any cracking and only minor faulting. The zero-maintenance 
slab thicknesses are 10.8 and 10.0 in. (274-254 mm) for JCP-34 and 35, 
respectively, with fatigue damage controlling the design. Dowel bars 
and a stabilized subbase are also recommended in this dry-freeze region. 
Thus, this recommended design should prevent both transverLd cracking and 
joint faulting and provide zero-maintenance over the 10 year period, and 
probably much longer. 

Other projects show similar results with the zero-maintenance slab 
design always exceeding the existing design. The long joint spacings on 
JCP-33 and 36 (30 and 25 ft. [9.1-7.6 m]) have a great effect on the 
zero-maintenance slab thickness. If joint spacing were reduced to 15 ft. 
(4.6 m) , for example, design slab thickness as required by fatigue is much 
less. Overall, while it is desirable to obtain additional data for further 
verification, the available results show that the new design procedure gives 



212 



designs that exceed existing designs that have provided long term 
maintenance-free performance. 



213 



CHAPTER 7 
CONCLUSIONS AND RECOMMENDATIONS 

7.1 CONCLUSIONS 

Comprehensive design procedures for heavily trafficked zero-maintenance 
plain jointed concrete pavements have been developed. This report (Vol. I) 
describes the field, laboratory and analytical studies upon which the 
procedures are based and provides research documentation. Based upon 
these results, a design manual was prepared (Vol. II) that contains all 
necessary procedures needed for actual design. A computer program 
designated JCP-1 was developed that is used to obtain fatigue damage and 
serviceability data for use in the structural design. The program is 
written in FORTRAN and is easily adaptable to most computers. The term 
"zero-maintenance" refers to the structural adequacy of the pavement 
lanes and shoulder. Thus, a "zero-maintenance" pavement would not 
require maintenance such as patching, joint repair, crack repair, grinding, 
and overlays. 

1. Field surveys of several plain jointed concrete pavements re- 
vealed several that have given zero-maintenance performance over time 
periods ranging up to 25 years under yery heavy traffic. Therefore, it 
is possible to achieve a "zero-maintenance" performance over long time 
periods. Several projects, however, exhibited distress that required 
maintenance. The following distress types commonly occurring in heavily 
trafficked plain jointed concrete pavements must be considered in design 
and thereby prevented: (1) joint faulting, (2) transverse cracking, 
(3) "D" cracking, (4) joint and corner spalling, (5) joint seal damage, 
(6) settlement of slabs, and (7) shoulder deterioration. 



214 



2. Comprehensive fatigue damage analysis procedures were developed 
that permit direct consideration of slab cracking. Stresses due to both 
traffic loading and thermal gradients are considered in the analysis 
through use of a finite element model. A fatigue damage limiting design 
criteria was determined from field data that provides a high reliability 
for preventing slab cracking. 

3. A new serviceability/performance design model was derived 
for plain jointed concrete pavements. The new model was derived 
from performance data of 25 sections of the original AASHO Road Test 

that have been under heavy mixed traffic on 1-80 since 1962. 
The performance model was "extended" using Westergaard' s edge stress model 
and through the incorporation of a climatic factor. The serviceability/ 
performance analysis provides consideration of various distress including 
joint faulting, slab cracking, settlement of slabs, spalling, etc. 

4. Design recommendations were developed for joint spacing, load 
transfer, and sealants. These factors were found to have a major effect 
on pavement zero-maintenance life. 

5. Use of full depth asphalt or concrete shoulders and subsurface 
drainage systems are strongly recommended. 

6. An example design application is provided that describes the 

use of the procedure in detail. The economic justification for constructing 

a zero-maintenance pavement is presented along with an example in Volume II. The 

additional cost increment for constructing a zero-maintenance pavement 

over a conventional pavement is determined for two different geographic 

areas and two levels of traffic and found to vary from 12 to 24 percent. 

7. Adequacy of the design procedures are assessed in terms of structural 

215 



sufficiency and also through a sensitivity analysis. The results show 
that the procedure provides designs that structurally exceed those of 
projects that have performed maintenance-free over long time periods 
and subject to heavy traffic. 

8. The design procedures and results documented herein can be used 
for numerous purposes other than zero-maintenance design. The effect of 
the following variables on slab cracking and performance can be analyzed: 
slab thickness, concrete strength and variation, foundation support 
(subbase and subgrade including degree of saturation), lane width, 
lateral distribution of traffic, thermal gradients, traffic overloads, 
lane distribution of trucks, joint spacing and others. 

7.2 RECOMMENDATIONS 

The zero-maintenance pavement design procedures documented herein 
and in the design manual (Ref. 1) are ready for trial implementation. 
They have been partially verified and shown to give adequate pavement 
structures. Many additional findings related to the design of plain 
jointed concrete pavements are believed to be significant and useful in 
minimizing the occurrence of distress and thus reducing maintenance 
costs. 

Trial implementation should proceed by selecting states that are willing 
to cooperate and are located in each of the four climatic regions. The state 
in consultation with FHWA and the project staff would select one or more 
heavily trafficked projects for consideration (either new designs or recon- 
structions). Zero-maintenance designs would then be developed by the state 
agency with the assistance of the project staff. The results would then be evalu- 
ated for adequacy by the state personnel, FHWA, and the Project staff. Any 
necessary modifications to the zero-maintenance design procedures be made so 
that it would be ready for implementation. 

216 





8 




9 




10 




11 



REFERENCES 



Darter, M. I., and E. J. Barenberg, "Design of Plain Jointed Concrete 
Pavements: Vol II - Design Manual," University of Illinois, Technical 
Report to Federal Highway Administration, June, 1977. 

Darter, M. I. and E. J. Barenberg, "Zero-Maintenance Pavement: Results 
of Field Studies on the Performance Requirements and Capabilities of 
Conventional Pavement Systems," Technical Report prepared for Federal 
Highway Administration, University of Illinois at Urbana-Champaign, 
April, 1976. 

Butler, B. C. , Jr., "Traffic Warrants for Premium Pavement Requiring 
Reduced Maintenance," Vol. I, II, III, D0T-FH-1 1 -8132, prepared for 
Federal Highway Administration, September, 1974. 

Treybig, H. J., Hudson, W. R. , and A. Abou-Ayyash, "Application of Slab 
Analysis Methods to Rigid Pavement Problems," Research Report 56-26, 
Center for Highway Research, The University of Texas at Austin, May, 1972. 

Eberhardt, A. C. and J. L. Wi Timer, "Computer Program for the Finite 
Element Analysis of Concrete Airfield Pavements," Technical Report S-26, 
Construction Engineering Research Laboratory, 1973. 

"The Multiple Wheel Load Elastic Layer Program," University of Illinois, 
Highway Pavements and Materials Group, Department of Civil Engineering. 

"The Finite Element Program for Pavement Analysis - User's Manual," 
Department of Civil Engineering, University of Illinois, 1973. 

Zienkiewicz, 0. C, and Y. K. Cheung, The Finite Element Method in Struc- 
tural and Continuum Mechanics , McGraw-Hill, 1967. 

"The AASHO Road Test - Report 5 - Pavement Research," Special Report 51 E, 
Highway Reserach Board, 1962. 

Hudson, W. R., and F. H. Scrivner, "AASHO Road Test Principal Relation- 
ships - Performance with Stress, Rigid Pavements," Special Report 73, 
Highway Research Board, 1962. 

Teller, L. W., and E. C. Southerland, "The Structural Design of Concrete 
Pavement," Part 1 - A Description of the Investgation, Public Roads, Vol. 
16, No. 8, 1935. 

12. Teller, L. W. , and E. C. Southerland, "The Structural Design of Concrete 
Pavements," Part 2 - Observed Effects of Variations in Temperature and 
Moisture on the Size, Shape and Stress Resistance of Concrete Pavement 
Slabs, Public Roads, Vol. 16, No. 9, 1935. 



217 



13. Teller, L. W., and E. C. Southerland, "The Structural Design of Concrete 
Pavements," Part 3 - A Study of Concrete Pavement Cross Sections, Public 
Roads, Vol. 16, No. 10, 1935. 

14. Teller, L. W. , and E. C. Southerland, "The Structural Design of Concrete 
Pavements," Part 4 - A Study of the Structural Action of Several Types of 
Transverse and Longitudinal Joint Designs, Public Roads, Vol. 17, Nos. 7 
and 8, 1936. 

15. Teller, L. W. , and E. C. Southerland, "The Structural Design of Concrete 
Pavements," Part 5 - An Experimental Study of the Westergaard Analysis of 
Stress Conditions in Concrete Pavement Slabs of Uniform Thickness, Public 
Roads, Vol. 23, No. 8, 1943. 

16. Hatt, W. K. , "Effect of Moisture on Concrete," Public Roads, Vol. 6, No. 1, 
March 1925; also Trans. ASCE, Vol. 84, 1926. 

17. Westergaard, H. M. , "Analysis of Stresses in Concrete Pavements Due to 
Variations in Temperature," Proc. Sixth Annual Meeting, Highway Research 
Board, 1927; also Public Roads, Vol. 8, No. 3, May, 1927. 

18. Bradbury, R. D. , Reinforced Concrete Pavements , Wire Reinforcement Insti- 
tute, Washington, D. C, 1938. 

19. Finney, E. A., and L. T. Oehler, "Final Report on Design Project, Michigan 
Test Road," Proc. Highway Research Board, Vol. 38, 1959. 

20. Carsberg, E. C. , and P. G. Velz, "Report on Experimental Project in 
Minnesota," Research Report 17-B, Highway Research Board, 1956. 

21. "Concrete Manual," Seventh Edition, United States Dept. of the Interior, 
Bureau of Reclamation, Denver, CO, 1966. 

22. Carlson, R. W. , "Drying Shrinkage of Concrete as Affected by Many Factors," 
Proceedings, Vol. 38, Part II, ASTM, 1938. 

23. Davis, R. E. , "A Summary of Investigations of Volume Changes in Cements, 
Mortars, and Concretes Produced by Causes Other Than Stress," Proceedings, 
Vol. 30, Part I, ASTM, 1930. 

24. Goldbeck, A. T. , "Friction Tests for Concrete on Various Subbases," ACI 
Proceedings, Vol. 8, 1917. 

25. Lin, T. Y., Design of Prestressed Concrete Structures , Wiley, Second 
Edition, New York, 1963. 

26. Goldbeck, A. T. , "Friction Tests of Concrete on Various Subbases," Public 
Roads, July, 1924. 

27. Harr, M. E., and G. A. Leonards, "Warping Stresses and Deflections in Con- 
crete Pavements," Proceedings, 38th Annual Meeting, 1959, Highway Research 
Board. 



218 



28. Wiseman, M. E. Harr, and G. A. Leonards, "Warping Stresses and Deflections 
in Concrete Pavements, Part II," presented at the 39th Annual Meeting, 1960, 
Highway Research Board. 

29. Moore, J. H., with discussions by E. C. Sutherland and Warner Harwood, 
"Thickness of Concrete Pavements," Transactions, ASCE, Vol. 121, Paper No. 
2834, 1956. 

30. Miner, M. A., "Cumulative Damage in Fatigue," Transactions, Am. Soc. of 
Mechanical Engr. , Vol. 67, 1945, pp. A159-A164. 

31. Kelley, E. F. , "Applications of the Results of Research to the Structural 
Design of Concrete Pavement," Public Roads, Vol. 20, No. 5, July, 1939. 

32. Murcock, J. W. , "A Critical Review of Research on Fatigue of Plain Concrete, 
Eng. Exp. Station, University of Illinois, Bulletin 476, 1965, 32 pp. 

33. Nordby, G. M. , "Fatigue of Concrete - A Review of Research," Proceedings, 
Am. Concrete Inst., Vol. 55, 1959, pp. 191-220. 

34. Statens institute for byggnadsforskning. Rapport 22: 1969: Utmattning 
av betond och armerad betong. En 1 itteraturoversi kt. (Fatigue of plain 
and reinforced concrete. A survey of literature), by B. Westerberg. 
Stockholm, Sweden, Svensk Byggtjanst, 1969. 

35. Hilsdorf, H. K. , and Kesler, C. E. , "Fatigue Strength of Concrete Under 
Varying Flexural Stresses," Proceedings, Am. Concrete Inst., Vol. 63, 
1966, pp. 1059-1976. 

36. Raithby, K. W. , and Whiffin, A. C, "Failure of Plain Concrete Under 
Fatigue Loading - A Review of Current Knowledge," Road Research Labora- 
tory, Report LR 231, Berkshire, England, 1968, 23 pp. 

37. McCall, J. T. , "Probability of Fatigue Failure of Plain Concrete," Pro- 
ceedings, Am. Concrete Inst., Vol. 55, 1959, pp. 233-245. 

38. Murdock, J. W., and Kesler, C. E. , "Effect of Range of Stress on Fatigue 
Strength of Plain Concrete Beams," Proceedings, Am. Concrete Inst., Vol. 
55,1959, pp. 221-231. 

39. Ople, F. S., Jr., and C. L. Hulsbos, "Probable Fatigue Life of Plain 
Concrete with Stress Gradient," Proceedings, Am. Concrete Inst., Vol. 
63, 1966, pp. 59-81. 

40. Clemmer, H. F. , "Fatigue of Concrete," Proceedings, Am. Soc. of Testing 
and Materials, Vol. 22, Part II, 1922, pp. 408-419. 

41. Kesler, C. E., "Effect of Speed of Testing on Flexural Strength of Plain 
Concrete," Proceedings, Highway Research Board, Vol. 32, 1953, pp. 251-258. 

42. Raithby, K. D. , and J. W. Galloway, "Effects of Moisture Condition, Age, 
and Rate of Loading on Fatigue of Plain Concrete, SP-41 , American Con- 
crete Institute, 1974. 

219 



43. Ballinger, C. A., "Cumulative Fatigue Damage Characteristics of Plain 
Concrete," Highway Research Record No. 370, Highway Research Board, 1972. 

44. Mills, . ., and . . Cawson. "Fatigue of Concrete," Highway Research 
Board Proceedings, Vol. , 1928. 

45. Darter, M. I., and E. J. Barenberg, "Zero-Maintenance Pavement: Results 
of Field'Studies on the Performance Requirements and Capabilities of 
Conventional Pavement Systems," Federal Highway Administration, Con- 
tract Number D0T-FH-1 1-8474, April 1976. 

46. "The AASHO Road Test - Pavement Research," Special Report 61 E, Highway 
Research Board, 1962. 

47. Hveem, F. N. , "Report on Experimental Project in California," Research 
Report 17-B, Highway Reserach Board, 1956. 

48. Coons, H. C. , "Report on Experimental Projection in Michigan," Research 
Report 17-B, Highway Research Board, 1956. 

49. "Thickness Design for Concrete Pavements," Portland Cement Association, 
1966. 

50. Carsberg, E. C. , and P. G. Velz, "Report on Experimental Project in 
Minnesota," Research Report 17-B, Highway Research Board, 1956. 

51. "Road Test One - MD," Special Report 4, Highway Reserach Board, 1952. 

52. Troxell, G. E., H. E. Davis, and J. W. Kelly, Composition and Properties 
of Concrete , 2nd Edition, McGraw-Hill, 1968. 

53. Taragin, Asriel, "Lateral Placement of Trucks on Two Lane Highways and 
Four-Lane Divided Highways," Public Roads, Vol. 30, No. 3, August, 1958. 

54. Emery, D. K. , "A Preliminary Report on the Transverse Lane Displacement 
for Design Trucks on Rural Freeways," paper presented to the ASCE Pavement 
Design Speciality Conference, Atlanta, GA, June, 1975. 

55. Dempsey, B. J., "A Heat Transfer Model for Evaluating Frost Action and 
Temperature-Related Effects in Multilayered Pavement Systems," Ph.D. Thesis, 
University of Illinois, Department of Civil Enginering, Urbana, Illinois, 
1969. 

56. Dempsey, B. J., and M. R. Thompson, "A Heat Transfer Model for Evaluating 
Frost Action and Temperature-Related Effects in Multilayered Pavement 
Systems, HRR NO. 342 , HRB, 1970, pp. 39-56. 

57. Van Breemen, W. , "Current Design of Concrete Pavements in New Jersey," 
HRB Proceedings, Vol. 28, 1948. 

58. "Structural Design of Portland Cement Concrete Pavements," Rev. Nov. 
1970, Illinois Department of Transportation. 



220 



59. "Structural Design of the Roadbed," April 16, 1970, Planning Manual, 
California Department of Transportation. 

60. AASHO Interim Guid for the Structural Desgin of Rigid Pavement 
Structures," Committee on Design, April, 1962. 

61. "AASHO Interim Guide for Design of Pavement Structures," Washington, 
D. C, 20004, 1972. 

62. McCullough, B. F. , et. al . , "Evaluation of AASHO Interim Guides for 
Design of Pavement Structures," NCHRP Report 128, Highway Research 
Board, 1972. 

63. "Combined Averaae Annual Precipitation and Evaporation MaD for the 
United States," compiled for the Federal Highway Administration, 
Robert G. Van Schooneveld, September, 1974. 

64. Corps of Engineers, "Engineering and Design, Pavement Design for Frost 
Conditions," EM-1110-345-306, 1958. 

65. Walker, R. S. and W. R. Hudson, "The Use of Spectral Estimates for 
Pavement Characterization," Research Report 156-2, CFHR, University 
of Texas, 1973. 

66. Huang, Y. H. , and S. T. Wang, "Finite Element Analysis of Concrete 
Slabs and Its Implications for Rigid Pavement Design," Highway Research 
Record 466, 1973. 

67. Huang, Y. H., "Finite Element Analysis of Slabs on Elastic Solids," 
Transportation Enqineerinq Journal of ASCE, Vol. 100, No. TE28, May 1974, 
pp. 403-416. 

68. "Computer Proqram for Concrete Airport Pavement Design," Special Report, 
Portland Cement Association, 1968. 

69. "Joint Design for Concrete Highway and Street Pavements," Portland 
Cement Association, 1975. 

70. "Joint Spacing in Concrete Pavements," Highway Research Board, Research 
Airport 17-B, 1956. 

71. Col ley, B. E. and H. A. Humphrey, "Aggregate Interlock at Joints in Con- 
crete Pavements," Highway Research Record No. 189, Highway Research 
Board, Washingtin, D. C, 1967. 

72. Stelzenmulter, W. B., L. L. Smith, and T. J. Larson, "Load Transfer at 
Contraction Joints in Plain Portland Cement Concrete Pavements," Research 
Report 90-D, Florida Department of Transportation, April, 1973. 

73. Darter, M. I., and W. B. Isakson, "Thermal Expansion and Contraction of 
Concrete Pavements in Utah," Interim Report, Project 915, Utah Department 
of Highways, 1970. 



221 



74. Gulden, Wouter, "Pavement Faulting Study," Final Report Project 7104, 
Georgia Department of Transportation, 1975. 

75. Spellman, D. L. , T. H. Woodstrom, and B. F. Neal , "Faulting of Portland 
Cement Concrete Pavements," Highway Research Record No. 407, Highway 
Research Board, 1972. 

76. Friberg, B. F. , "Design of Dowels in Transverse Joints of Concrete 
Pavements," Transactions, ASCE, Vol. 105, 1940. 

77. "Design, Construction, and Maintenance of PCC Pavement Joints," Synthe- 
sis of Highway Practice No. 19, NCHRP, Washington, D. C, 1973. 

78. "Evaluation of the Performance of Doweled Contraction Joints Placed on 
Three Types of Subbase Courses," Report No. 30, unpublished internal 
reports, Georgia Department of Transportation, February, 1974. 

79. Griffin, H. W. , "Transverse Joints in the Design of Heavy Duty Concrete 
Pavement," Proceedings HRB, Vol. 23, 1943. 

80. Van Breeman, W. , "Experimental Dowel Installations in New Jersey," 
Proc. HRB, Vol. 34, 1955. 

81. Van Breeman, W. , "Special Papers on the Pumping Action of Concrete 
Pavements," Research Reports No. ID, Highway Research Board, 1945. 

82. Cedergren, H. R. , Drainage of Highway and Airfield Pavements , John Wiley 
& Sons, New York, 1974. 

83. "Implementation Package for a Drainage Blanket in Highway Pavement 
Systems," Federal Highway Administration, May, 1972. 

84. Barksdale, R. D., and R. G. Hicks, "Improved Pavement Shoulder Joint 
Design," Final Report NCHRP Project 14-3, June, 1975. 

85. Cedergren, H. R. , et al., "Guidelines for the Design of Subsurface 
Drainage Systems for Highway Structural Sections," Technical Report 
prepared for Federal Highway Administration, June, 1972. 

86. Cedergren, H. R. , "Methodology and Effectiveness of Drainage System for 
Airfield Pavements," Technical Report E-13, CERL, 1974. 

87. Stromberg, F. J., and J. Weisner, "Inservice Behavior of Preformed Neo- 
prene Joint Seals," Interim Report, Maryland State Highway Administration, 
1972. 

88. McBride, J. C, and M. S. Decker, "Performance Evaluation of Utah's 
Concrete Pavement Joint Seals," Final Report, Utah Department of Trans- 
portation, October, 1974. 

89. Carlson, R. D., "Transverse Joint Construction and Sealing Practices," 
Research Report 20, New York State Department of Transportation, 1974. 



222 



90. "Evaluation of Preformed Elastomeric Pavement Joint Sealing Systems 
and Practices," NCHRP Research Results Digest 35, 1972. 

91. Darter, M. I. and M. L. Hine, "PCC Joint Dowel Pullout Test Results," 
unpublished Report, Department of Civil Engineering, University of 
Illinois at Urbana-Champaign, 1975. 

92. Bryden, J. E. and R. G. Phillips, "Performance of Transverse Joint 
Supports in Rigid Pavements," Research Report 12, New York State 
Department of Transportation, March, 1973. 

93. Bryden, J. E., and R. G. Phillips, "New York's Experience with Plastic- 
Coated Dowels," Special Report 27, New York Department of Transportation, 
December, 1974. 

94. "Structural Design Considerations for Pavement Joints," ACI Committee 
325, ACI Journal, July, 1956. 

95. Barksdale, R. D. , and R. G. Hicks, "Improved Pavement Shoulder Joint 
Design," Final Report NCHRP Project 14-3, June, 1975, 151 pps. 

96. Portigo, Josette M. , "State of the Art Review of Paved Shoulders," 
paper presented at the 55th Annual Meeting of the Transportation 
Research Board, Washington, D. C, January, 1976. 

97. Report of an Investigation on Potential Causes of Deterioration of the 
Shoulders on the Southwest Expressway, Professor E. J. Barenberg and Pro- 
fessor M. R. Thompson, Civil Engineering Department, University of 
Illinois, October 12, 1965. 

98. Report No. 27, Portland Cement Concrete Shoulders (IHR-404), Illinois 
Divisions of Highways in cooperation with the U. S. Department of 
Transportation, Federal Highway Administration, and Bureau of Public 
Roads, July, 1970. 

99. McKenzie, Lloyd J., Report No. 39, Experimental Paved Shoulders on 
Frost Susceptible Soils (IHR-404), Illinois DOT in cooperation with 
U. S. DOT and Federal Highway Administration. 

100. "Concrete Shoulders." Portland Cement Association, 1975. 

101. Spellman, D. L., J. H. Woodstrom, B. F. Neal , and P. E. Mason, 
"Recent Experimental PCC Pavements in California," Interim Report, 
California Division of Highways, Nov. 1072. 



223 



APPENDIX A 

A.l INPUT GUIDE - JCP-1 PROGRAM 
ZERO-MAINTENANCE DESIGN OF PLAIN JOINTED CONCRETE PAVEMENT 

IDENTIFICATION OF PROBLEM 



Three Cards 



20A4 



20A4 



20A4 



80 



Enter descriptive identification of design project; date of run, project 
number, designer, etc. (Any or all of the cards may be left blank). 

DESIGN CRITERIA DATA 

One Card 



F10.0 


F10.0 


F10.0 


F10.0 


15 





1 10 20 30 40 45 
DLIFE SIC PT OPEN KMONTH 

DLIFE = Pavement zero-maintenance design life (years) 

SIC = Initial serviceability index after construction 

PT = Terminal serviceability index for zero-maintenance 

OPEN = Time after PCC slab placement that pavement will be 
opened to traffic (years) 

KMONTH = Month pavement will be opened to traffic (right justify) 
(Input 1 through 12 according to the following key: 
Jan=l, Feb=2, Mar=3, Apr=4, May=5, Jun=6, Jul=7 
Aug=8, Sep=9, 0ct=10, Nov=ll, Dec=12) 

PRINTOUT DATA CONTROL 

One Card 



80 



8011 



DLIFE 

Enter 1 in the columns that correspond to the years during which summary 
of fatigue and serviceability data will be printed. 



80 



One Card 



8011 



DLIFE 



80 



Enter 1 in the columns that correspond to the years during which 
comprehensive fatigue output will be printed. 

SLAB PROPERTIES DATA 

One Card 



F5.0 



F5.0 



F5.0 



F5.0 



E10.3 



E10.3 



15 10 15 20 
H L FF FCV ET 



30 



40 



80 



H = Slab thickness - inches 

L = Slab length - feet 

FF = Mean PCC modulus of rupture (28 days) - psi 

FCV = Coefficient of variation of PCC modulus of rupture - percent 

ET = PCC coefficient of thermal expansion (per degree - F) 

E = PCC modulus of elasticity (psi) 

TRAFFIC DATA 
One Card 



F10.0 


F10.0 


F10.0 


F10.0 


F10.0 


F10.0 


F10.0 


15 





1 



10 



20 



30 



40 



50 



60 



70 75 80 
D 






ADTI ADTF T LD DD A PC 

ADTI = Average daily traffic at beginning of design period - two direction 
ADTF = Average daily traffic at end of design period - two direction 
T = Percent trucks of ADT 

LD = Percent trucks in heaviest traveled or design lane 
DD = Percent direction distribution 



225 



SIX CARDS SET FOR EACH ADDITIONAL TRIAL PCC SLAB THICKNESS 
IDENTIFICATION OF PROBLEM 
Three Cards (Same as first trial thickness). 



20A4 



20A4 



20A4 



80 



SLAB THICKNESS 



One Card 



F5.0 



1 5 

H 



80 



H = New Slab Thickness, inches 



CONTROL DATA 
One Card 



8011 



DLIFE 

Enter 1 in the columns that correspond to the years during which summary 
of fatigue and serviceability data will be printed. 

One Card 



80 



8011 



DLIFE 80 

Enter 1 in the columns that correspond to the years during which 
comprehensive fatigue output will be printed. 

FINAL CARD OF DATA DECK 



12 



80 



/* indicates end of data deck 



226 



A = Mean axles per truck 

PC = Percent trucks during daylight 

D = Mean distance from slab edge to outside of truck duals (in.) 
(right justify) 

One Card 



15 



15 



1 5 10 
KK KSAL 



80 



KK = Number of axle load distribution groups (single plus tandem) 
(right justify) 

KSAL = Number of single axle load distribution groups (right justify) 



As Many Cards As Needed 



F10.0 


F10.0 


F10.0 


F10.0 


F10.0 


F10.0 


F10.0 


F10.0 


F10.0 


F10.0 


F10.0 


F10.0 


F10.0 


F10.0 


F10.0 


F10.0 



1 10 20 30 40 50 60 70 
LOAD(I) 

[LOAD(I), 1=1, KK] 

LOAD(I) = The highest value of each axle load distribution group 
(first enter single axle loads (KSAL) and then tandem 
axle loads)(pounds) 

As Many Cards As Needed 



80 



F10.0 


F10.0 


F10.0 


F10.0 


F10.0 


F10.0 


F10.0 


F10.0 


F10.0 


F10.0 


F10.0 


F10.0 


F10.0 


F10.0 


F10.0 


F10.0 



1 10 
DIST(I) 

[DIST(I),I=1,KK] 



20 



30 



40 



50 



60 



70 



DIST(I) = The percentage axle loads in each of the KK axle load groups 
input in the previous card (first enter single axle percentage 
and then tandem axle percentage) 



80 



227 



One Card: 



F5.0 



F5.0 



F5.0 



F5.0 



F5.0 



F5.0 



F5.0 



F5.0 



F5.0 



F5.0 



F5.0 



F5.0 



15 10 15 20 25 30 35 40 45 50 55 60 80 

TRUKPC(I) 

[TRUKPC(I),I=1,12] 

TRUKPC(I) = The monthly truck percentage over year (enter percentage for 
first month pavement will be opened to traffic in Columns 1-5, 
2nd month in Columns 6-10, etc.) 

FOUNDATION SUPPORT DATA : 
One Card: 



E 



3 



F5.0 



F5.0 



F5.0 F5.0 



F5.0 



F5 



5.0 



F5.0 



F5.0 



F5.0 



F5.0 



F5.0 



1 5 
K(J) 



10 15 20 25 30 



35 



40 45 50 55 



60 



80 



[K(J),J=1,12] 

K(J) = Modulus of foundation support (k-value at top of subbase) for 
each month in pci (enter k-value for first month pavement will 
be opened to traffic in Columns 1-5, 2nd month in Columns 6-10, etc.) 

One Card: 



F10.0 



1 10 
ERODEF 

ERODEF 



80 



The amount of erodability of foundation at the end of design life 
in inches. 



One Card 



F10.0 



1 10 
DK 



80 



DK = Design modulus of foundation support for serviceability/performance 
analysis in pci . 



228 





F10.0 


F10.0 


F10.0 


F10.0 


F10.0 


F10.0 





10 20 30 40 50 60 70 80 
G2(l,2) G2(2,2) G2(3,2) G2(4,2) G2(5,2) G2(6,2) 





fio.o ho.o 


F10.0 


FIO.O 


FIO.O 


FIO.O 





1 10 20 30 40 50 60 70 80 
G2(7,2) G2(8,2) G2(9,2) G2(10,2) G2(ll,2) G2(12,2) 

H2 = PCC slab thickness for relatively thick slab (usually 12 inches) - inches 

G2(J,M) = Mean temperature gradients for slab of H2 thickness for each month, 
day, and night where 

J = index for months 

(J=l for first month opened to traffic) 
M = index for day (NM) and night (M=2) 

One Card 



FIO.O 



1 10 
RF 



80 



RF = Regional factor 

The cards shown on the following page may be added for each additional 
PCC slab thickness to be analyzed; 



229 



ENVIRONMENTAL DATA (TWO SETS OF FOUR CARDS) 
Set One - Four Cards 



F10.0 


F10.0 


F10.0 


F10.0 


F10.0 


F10.0 


F10.0 





1 10 20 30 40 50 60 70 

HI Gl(l,l) Gl(2,l) Gl(3,l) Gl(4,l) Gl(5,l) Gl(6,l) 



80 



F10.0 



F10.0 F10.0 



F10.0 



F10.0 



F10.0 



10 20 30 40 50 60 70 

Gl(7,l) Gl(8,l) Gl(9,l) Gl (10,1 ) Gl(ll,l) Gl(12,l) 



80 





F10.0 


F10.0 


F10.0 


F10.0 


F10.0 


F10.0 



10 20 30 40 50 60 70 

Gl(l,2) Gl (2,2) Gl (3,2) Gl (4,2) Gl (5,2) Gl (6,2) 



80 





F10.0 


F10.0 


F10.0 


F10.0 


F10.0 


F10.0 



1 10 20 30 40 50 60 70 80 
Gl (7,2) Gl (8,2) Gl (9,2) Gl (10,2) Gl(ll,2) Gl(12,2) 

HI = PCC slab thickness for relatively thin slab (usually 8 inches) - inches 

G1(J,M) = Mean temperature gradients for slab of HI thickness for each month, 
day, and night where 

J = index for months 

(J=l for first month opened to traffic) 
M = index for day (M=l) and night (M=2) 



Set Two - Four Cards 



F10.0 


F10.0 


F10.0 


F10.0 


F10.0 


F10.0 


F10.0 





1 10 20 30 40 50 60 70 80 
H2 G2(l,l) G2(2,l) G2(3,l) G2(4,l) G2(5,l) G2(6,l) 



F10.0 



F10.0 



F10.0 



F10.0 



F10.0 



F10.0 



10 20 30 40 50 60 70 80 
G2(7,l) G2(8,l) G2(9,l) G2(10,l) G2(ll,l) G2(12,l) 



230 



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232 



A. 3 FLOW CHART OF PROGRAM 




START 



D 



READ 
INPUTS 



PRINT 
INPUTS 



I 



MAKE SEVERAL 
CALCULATIONS 
FOR TRAFFIC, 
PCC, ERODA- 
ILITY, LATERAL 
DISTRIBUTION 




233 



t 




CALCULATE FATIGUE 
DAMAGE FOR NY YEAR, 
J MONTH, I DAY/NIGHT 
AND KK AXLE LOAD 




234 



t 



* I 



PRINT 
MONTHLY/YEARLY 
FATIGUE DAMAGE 



CALCULATE 
YEARLY SERVICE 
ABILITY AND 
18-KIP ESAL 



PRINT 
SERVICEABILITY AND 
18-KIP ESAL 




235 




i 



PRINT 

SUMMARY OF RESULTS 

AT END OF DESIGN 
PERIOD 



DO NEXT PROBLEM 
IF REQUIRED 




STOP 



236 



A. 4 ..1CP-1 PROGRAM LISTING 



# » 
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253 



ftU.S. GOVERNMENT PRINTING OFFICE: 1978 O — 621-625/268 REGION 3-1 



FEDERALLY COORDINATED PROGRAM OF HIGHWAY 
RESEARCH AND DEVELOPMENT (FCP) 



The Offices of Research and Development of the 
Federal Highway Administration are responsible 
for a broad program of research with resources 
including its own staff, contract programs, and a 
Federal-Aid program which is conducted by or 
through the State highway departments and which 
also finances the National Cooperative Highway 
Research Program managed by the Transportation 
Research Board. The Federally Coordinated Pro- 
gram of Highway Research and Development 
(FCP) is a carefully selected group of projects 
aimed at urgent, national problems, which concen- 
trates these resources on these problems to obtain 
timelv solutions. Virtually all of the available 
funds and staff resources are a part of the FCP. 
together with as much of the Federal-aid research 
funds of the States and the NCHRP resources as 
the States asree to devote to these projects." 



FCP Category Descriptions 

1. Improved Highway Design and Opera- 
tion for Safety 

Safetv R&D addresses problems connected with 
the responsibilities of the Federal Highway 
Administration under the Highway Safety Act 
and includes investigation of appropriate design 
standards, roadside hardware, signing, and 
phvsical and scientific data for the formulation 
of improved safety regulations. 

2. Reduction of Traffic Congestion and 
Improved Operational Efficiency 

Traffic R&D is concerned with increasing the 
operational efficiency of existing highways by 
advancing technologv. by improving designs for 
existing as well as new facilities, and by keep- 
ing the demand-capacity relationship in better 
balance through traffic management techniques 
such as bus and carpool preferential treatment, 
motorist information, and rerouting of traffic. 



* The complete T-volume official statement of the FCP is 
available from the National Technical Information Service 
(XTISl. Springfield. Virginia 22161 (Order No. PB 242057. 
price $45 postpaid). Single copies of the introductory 
volume are obtainable without charge from Program 
Analysis (HRD-2). Offices of Research and Development, 
Federal Highwav Administration. Washington, D.C. 20500. 



3. Environmental Considerations in High- 
way Design, Location, Construction, and 
Operation 

Environmental R&D is directed toward identifv- 
ing and evaluating highway elements which 
affect the quality* of the human environment. 
The ultimate goals are reduction of adverse high- 
way and traffic impacts, and protection and 
enhancement of the environment. 

4. Improved Materials Utilization and Dura- 
bility 

Materials R&D is concerned with expanding the 
knowledge of materials properties and technologv 
to fullv utilize available naturally occurring 
materials, to develop extender or substitute ma- 
terials for materials in short supplv. and to 
devise procedures for converting industrial and 
other wastes into useful highwav products. 
These activities are all directed toward the com- 
mon goals of lowering the cost of highwav 
construction and extending the period of main- 
tenance-free operation. 

5. Improved Design to Reduce Costs, Extend 
Life Expectancy, and Insure Structural 
Safety 

Structural R&D is concerned with furthering the 
latest technological advances in structural de- 
signs, fabrication processes, and construction 
techniques, to provide safe, efficient highways 
at reasonable cost. 

6. Prototype Development and Implementa- 
tion of Research 

This category is concerned with developing and 
transferring research and technology into prac- 
tice, or. as it has been commonlv identified, 
"technology transfer." 

7. Improved Technology for Highway Main- 
tenance 

Maintenance R&D objectives include the develop- 
ment and application of new technology to im- 
prove management, to augment the utilization 
of resources, and to increase operational efficiency 
and safetv in the maintenance of highway 
facilities. 



DOT UBRMN 

Jill 




R&D